Regenerative Fuel Cell Systems - ACS Publications - American

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Energy & Fuels 1998, 12, 56-71

Regenerative Fuel Cell Systems Fred Mitlitsky,* Blake Myers, and Andrew H. Weisberg Lawrence Livermore National Laboratory, 7000 East Avenue, L-045, Livermore, California 94551 Received August 22, 1997. Revised Manuscript Received October 2, 1997X

Regenerative fuel cell (RFC) systems produce power and electrolytically regenerate their reactants using stacks of electrochemical cells. Energy storage systems with extremely high specific energy (>400 Wh/kg) have been designed that use lightweight pressure vessels to contain the gases generated by reversible (unitized) regenerative fuel cells (URFCs). Progress is reported on the development, integration, and operation of rechargeable energy storage systems with such high specific energy. A primary fuel cell test rig with a single cell (46 cm2 active area) has been modified and operated reversibly as a URFC (for up to 2010 cycles on a single cell). This URFC uses bifunctional electrodes (oxidation and reduction electrodes reverse roles when switching from charge to discharge, as with a rechargeable battery) and cathode feed electrolysis (water is fed from the hydrogen side of the cell). Lightweight pressure vessels with performance factors (burst pressure × internal volume/tank weight) > 50 km (2.0 million in.) have been designed, and a vessel with performance factor of 40 km (1.6 million in.) was fabricated. These vessels use lightweight bladder liners that act as inflatable mandrels for composite overwrap and provide the permeation barrier for gas storage. Bladders are fabricated using materials that are compatible with humidified, electrolyzed gases and are designed to be compatible with elevated temperatures that occur during fast fills or epoxy curing cycles. RFC systems are considered that use hydrogen/oxygen, hydrogen/air, or hydrogen/halogen chemistries. Hydrogen/halogen URFCs are capable of higher round-trip efficiency than hydrogen/oxygen URFCs but are significantly heavier. Therefore, hydrogen/halogen URFCs are advantageous for stationary applications, whereas hydrogen/oxygen URFCs are advantageous for mobile applications. Safety aspects of halogens may prohibit their use in densely populated areas and some commercial applications, so these niches can also benefit from hydrogen/oxygen URFCs.

Introduction NASA Dryden Flight Research Center is supporting research and development on regenerative fuel cells (RFCs) as an enabling technology for high-altitude long endurance (HALE) solar rechargeable aircraft (SRA), under its Environmental Research Aircraft and Sensors Technology (ERAST) program. This research is being leveraged by the U.S. Department of Energy to develop similar systems for ground transportation and utility applications. A significant advance in energy storage is being developed to address a range of applications including HALE SRA, zero emission vehicles (ZEVs), hybrid energy storage/propulsion systems for spacecraft, energy storage for remote (off-grid) power sources, and peak shaving for on-grid applications.1-5 Innovative energy storage designs are being implemented, and * Corresponding author. Phone: (510) 423-4852. Fax: (510) 4239178. E-mail [email protected]. X Abstract published in Advance ACS Abstracts, December 1, 1997. (1) de Groot, W. A.; Arrington, L. A.; McElroy, J. F.; Mitlitsky, F.; Weisberg, A. H.; Carter II, P. H.; Myers, B.; Reed, B. D. 33rd AIAA/ ASME/SAE/ASEE Joint Propulsion Conf., July 6-9, 1997 1997, AIAA 97-2948. (2) Mitlitsky, F.; Myers, B.; Weisberg, A. H. 1996 Fuel Cell Seminar 1996, UCRL-JC-125220, 743-746. (3) Mitlitsky, F.; de Groot, W.; Butler, L.; McElroy, J. F. AIAA Small Satellite Conf., Sept 16-20, 1996 1996, UCRL-JC-125242. (4) Mitlitsky, F.; Colella, N. J.; Myers, B. 1994 Fuel Cell Seminar 1994, UCRL-JC-117130, 624-627.

their preliminary testing already demonstrates the capability to significantly outperform foreseeable battery technologies. One performance measure these developments are designed to excel at is specific energy, which describes how much energy can be stored per unit mass (Wh/kg or J/kg). Designs capable of delivering a significant advance in specific energy must openly assume innovations in key components, and integration experience with advanced technologies that have yet to be operated in combination. Investigation of the performance and operating characteristics of the two classes of critical components, electrochemical power converters (RFCs) and vessels capable of containing pressurized reactants, is proceeding independently toward implementations likely to minimize system mass. Lightweight pressure vessels with high cycle life and low permeability to stored reactants are the other class of critical components that call for innovative development to enable RFC systems that excel at lightweight energy storage. Encouraging test results reported herein demonstrate that both classes of innovative components can already provide the performance levels necessary to realize unprecedented specific energy storage levels (above 400 (5) Mitlitsky, F.; Colella, N. J.; Myers, B.; Anderson, C. J. 28th IECEC 1993, 1, 1.1255-1.1262; UCRL-JC-113485.

S0887-0624(97)00151-5 CCC: $15.00 © 1998 American Chemical Society Published on Web 01/12/1998

Regenerative Fuel Cell Systems

Energy & Fuels, Vol. 12, No. 1, 1998 57

Figure 1. Water rocket block diagram.

Wh/kg) in the near term. Further integration testing is planned to prove the performance of entire energy storage systems over the next few years. Preliminary integration results emerging from component tests are also reported that prove the most mass-conserving electrochemical power converters, reversible (unitized) regenerative fuel cells (URFCs), have adequate cycle life to build useful energy storage systems. Limitations of Batteries. The search for a nearterm, flightweight solution to propelling a solar-powered aircraft at night on stored solar energy created the original rationale for this cluster of innovations. Rechargeable batteries provide the proven route to electrical energy storage, but even the projected types are far too heavy for acceptable HALE SRA designs. Considerable spreadsheet analysis of aircraft system requirements endorsed the development of URFCs, coupled with novel lightweight pressure vessels that are integrated into the structure of the aircraft. Such systems were designed to have exceedingly high specific energies (energy/weight) of >400 Wh/kg.4 Related terrestrial applications including portable power supplies, automobiles, and remote power sources were investigated with similar top-down design tools, and were found to require less demanding subsets of the same development plan. Integrating the best components of the SRA energy storage technology with nontoxic propulsion, and sharing function with structure mass already required in spacecraft, are expected to offer an extremely attractive approach for spacecraft applications.1,3 Figure 1 shows a schematic of the components interconnected in such a spacecraft. All of the applications discussed herein can be visualized as subsets of the Figure 1 schematic, although real implementations will also include various ancillary components (valves, pumps, regulators, tubing, cabling, insulation, connectors, and controllers). For spaceborne applications, as well as for HALE aircraft, lightweight energy storage would best be achieved by

Figure 2. Pathfinder solar-powered aircraft.

coupling URFCs with lightweight pressure vessels that are integrated into the vehicle’s structure. Hydrogen and oxygen that are made by the solar-powered electrolysis of water can be used for energy storage, or as propellants for a high specific impulse, high thrust-toweight ratio rocket propulsion system. The nontoxic nature of water, and the ability to convert it into rocket propellants on the fly, may provide the strongest incentives to develop spaceborne applications, which also pose the toughest development challenges. HALE SRA requirements demand specific energy levels that are within the capabilities of URFC and pressure vessel components. The Pathfinder solarpowered aircraft is a flying wing that was designed and manufactured, and is operated, by AeroVironment, Inc., of Simi Valley, CA. Pathfinder measures 99 ft (30 m) wingspan and 8 ft (2.4 m) chord. The total weight for this ∼800 ft2 (74 m2) aircraft is ∼500 lb (230 kg). Pathfinder set a solar-powered altitude record of 50 500 ft (15.4 km) on September 11, 1995 (see Figure 2). The plane took off and landed at the NASA Dryden Research

58 Energy & Fuels, Vol. 12, No. 1, 1998

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Table 1. Specific Energy of URFC and Rechargeable Batteries battery system

theoretical specific energy (Wh/kg)

packaged specific energy (Wh/kg)

comments

H2/O2 URFC Li-SPE/MOx Ag/Zn Li/LiCoO2 Li/AlFeS2 Na/S Li/TiS2 Li/ion Ni/Zn Ni/MHx Ni/H2 Ni/Cd Pb/acid

3660 735 450 735 515 1180 470 700 305 470 470 240 170

400-1000 220 200 150 150 150 130 100 (135a) 90 70 (85a) 60 60 50

URFC with lightweight pressure vessels Li-solid polymer electrolyte/ metal oxide, novel packaging excess Zn required for high cycle life, low charge rate poor cycle life, high capacity fade g400 °C thermal management ∼350 °C thermal management ∼50% DOD for high cycle life (900 cycles) projection revised Nov 1996 excess Zn required, low specific energy MHx is metal hydride, projection revised Nov 1996 low specific energy low specific energy low specific energy

a

Projections revised in November 1996 by private communications with B. M. Barnett (A. D. Little, Inc.).

Center, in Edwards, CA, and was solar powered during take off, climb, and most of the descent. Primary lithium batteries were used to maneuver and controllably land the plane at night. During the latest series of flight tests (which are ongoing during the writing of this paper) at the Pacific Missile Range Facility, Kauai, Hawaii, Pathfinder set a new world record of 71 500 ft (21.8 km) on July 7, 1997. In addition to exceeding the solar-powered altitude record, this recent flight set the record for the highest altitude ever attained by any propeller-driven aircraft. The previous record of 67 028 ft (20.43 km) was held by the experimental Boeing Condor, which was powered by a multistage turbocharged engine burning fossil fuel.6 Photographs of Pathfinder are available on the World Wide Web site set up by NASA Dryden at http://www.dfrc.nasa.gov/ PhotoServer/Pathfinder/index.html. One of the goals of the NASA ERAST program is to design and build a plane that has sufficient energy storage on board to enable level flight at high altitudes for multiple days or weeks. A particular design to accomplish this would have ∼220 ft (67 m) wingspan and assumes the use of RFCs with lightweight tanks integrated into the main structural wingspar. The conceptual arrangement for such a system is a subset of the spacecraft schematic shown in Figure 1, without the thrusters. An airborne system can make even better use of the integration of pressure vessels for reactant gas storage with a vehicle’s structure, especially in a flying wing configuration, where almost all of the vehicle’s structure mass resides in the wingspar. Because that wingspar must be designed to withstand high bending loads, much of the structure mass requirements that would make pressure vessels burdensome are already supplied by the spar’s structure. Energy storage requirements were investigated for the design of the next-generation HALE SRA. Specific energy >400 Wh/kg and efficiency ∼43-50% will be required for many desired missions. Rechargeable batteries (see Table 1) and flywheels are projected to have insufficient packaged specific energy to achieve this goal. The difference between “theoretical” and “packaged” specific energies of battery technologies is highlighted in Figure 3. Theoretical specific energy uses the weight of stoichiometric reactants only. Optimum specific energy could be achieved with stoichiometric (6) NASA Headquarters Press Release 97-153, July 14, 1997; http: //www.dfrc.nasa.gov/PAO/PressReleases/1997/97-153.html.

Figure 3. Theoretical vs packaged specific energy.

complete cells, which add the weight of electrolyte, separators, and current collectors required to draw current from the cell. Functional cells add the weight of nonstoichiometric reactants necessary to achieve the application’s required cycle life. Packaged cells add the weight of packaging required for safe containment, shipping, handling, and use in real systems. The packaged specific energy is based on packaged cells and typically does not achieve better than ∼30% of the theoretical specific energy, especially when high cycle life is required. Estimates for the packaged specific energies of batteries in Table 1 were performed by A. D. Little, Inc. under contract to Lawrence Livermore National Laboratory (LLNL) in 1993, and assume novel packaging for large systems.5 Since then, there have been some advances in packaged specific energies projected for lithium ion (∼135 Wh/kg) and nickel metal hydride (∼85 Wh/kg). These improvements do not change the conclusion that rechargeable batteries are not capable of achieving the required >400 Wh/kg. The estimate for URFCs and lightweight tankage assumes several component and integration innovations that are currently being investigated. Energy Storage Designs. RFC-based energy storage alternatives free the vehicle designer from many of the hidden compromises that accompany battery technologies. RFC energy storage is uncoupled from power generation or consumption, because reactants are stored outside any electrochemical cells, in reactant tanks that can be sized independently. Power ratings determine the cell stack size, but not the sizing of the containers that hold reactants. Energy storage requirements size the capacity of reactant containers independently of the RFC stack(s). More tank capacity can be added to store more energy, without redesigning electrodes. Since reactants are actively transported to/from the RFC cell stack, depth of discharge does not reduce available power. Many RFC applications require tanks that are


Energy & Fuels, Vol. 12, No. 1, 1998 59 Table 2. Specific Energy of RFC Designs

stack type

stack wt (kg)

ancillary wt (kg)

tank wt (kg)

fuel wt (kg)

layout penalty wt (kg)

system wt (kg)

specific energy (Wh/kg)

comments

1 URFC 2 URFCs 3 URFCs 1 FC/1 EC 2 FCs/1 EC 3 FCs/3 ECs

41.8 49.6 55.0 63.2 79.1 73.2

16.4 26.4 36.4 28.6 38.6 73.2

26.8 26.8 26.8 26.8 26.8 26.8

32.3 32.3 32.3 32.3 32.3 32.3

12.7 0.0 0.9 38.6 44.1 17.7

130 135 151 190 221 258

431 415 370 295 253 217

not redundant lightweight redundant, favored redundant, heavy not redundant, heavy partly redundant, heavy redundant, heavy

heavier than their cell stacks, so that their system mass may be dominated by tankage mass for even the lightest weight tank designs discussed below. Applications where energy storage times (useful energy capacity/ rated power; or specific energy/specific power; or typical discharge times) are longer than ∼1 h will generally lead to the case where tankage dominates energy storage mass. If heavy tanks are used, this energy storage time can be as short as several minutes. RFC system designs require lightweight gas storage to achieve the 400 Wh/kg levels necessary for HALE SRA missions. A lightweight gas storage system was designed and prototyped at LLNL that uses bladders integrated into the wingspar to minimize tank mass. This bladder design is ∼100 lb (45 kg) lighter (∼500 kg gross weight) than similar tanks with other liners. Hydrogen permeation losses should be adequately low (400 Wh/kg for a fully packaged energy storage system (based on one or two URFCs) depending upon lightweight pressure vessels that are integrated into the wingspar, to form a structural tank. Structural Tanks. The integration of structure and tankage requirements is particularly promising for spaceborne energy storage applications, where mass and volume are most severely constrained. Structural tanks can range from applications which are predominantly structure (e.g., the wingspar of HALE SRA) to applications which are predominantly tankage (e.g., tankage for gaseous fueled automobiles). The importance of lightweight pressurized gas containment for other types of vehicles, and many alternative energy technologies that employ hydrogen, have prompted a more broadly applicable development of extremely lightweight pressure vessel technology at LLNL that will be discussed below. The strength per weight of a tubular spar (assuming constant mass per unit length of a given material) increases linearly with tube diameter and is not sensitive to wall thickness. Buckling resistance per weight of a tubular spar (assuming constant mass per unit

Table 3. Wingspar Loading Conditions loading condition

maximum value

+5g flight shear +5g flight moment +5g flight torsion control deflection torsion landing shear landing moment landing torsion gas storage pressure

14500 lb (64500 N) 29000 ft‚lb (39300 N‚m) 240 ft‚lb (325 N‚m) 760 ft‚lb (1030 N‚m) 10000 lb (44000 N) 12200 ft‚lb (16500 N‚m) 280 ft‚lb (380 N‚m) 300 psia (2.1 MPa)

length of a given material) increases linearly with wall thickness and is not directly dependent on tube diameter. The strength per weight and buckling resistance per weight criteria lead to an optimum tube diameter and wall thickness for fabricating a lightweight structure using a given material. For HALE SRA designs, tube diameter is constrained by airfoil geometry. A tubular spar can provide a large volume for storing gaseous reactants if it is transformed into a structural tank by adding some additional composite overwrap, bulkheads, and a permeation barrier. Larger diameter tubes increase the volume available for reactant storage and will reduce the peak pressure required for fixed capacity. Reducing the peak pressure requirements results in improved electrolyzer efficiency (and roundtrip efficiency), but can make the tubes buckling-critical rather than strength-critical. Table 3 summarizes preliminary wingspar structural requirements and system tankage requirements assuming that tankage is integrated into a tubular wingspar for most of an aircraft’s wingspan.5 The stresses induced by the large flight moment in this design are significantly larger than the stresses induced by the relatively low pressure loads. As a result, the tank mass can represent a relatively small fraction of the composite material required for structure, and its presence will thicken the spar wall and help improve buckling resistance of the large diameter structure. This design requires a lightweight permeation barrier to enable gas storage within the volume of the composite wingspar. Permeation Barriers. Various materials were investigated for their potential to serve as hydrogen permeation barriers inside composite pressure vessels. Table 4 is an excerpt of the best barrier candidates from a tabulation of the hydrogen permeabilities (K) for thick samples of various materials.7 Hydrogen permeability for a given temperature can be calculated from the table by using the following formulas:

K ) K0 exp(-θK/T) for metals K ) K0T exp(-θK/T) for glasses The table entries are ranked in terms of extrapolated permeabilities at room temperature, with the best (7) Souers, P. C. Hydrogen Properties for Fusion Energy; University of California Press: Berkeley, CA, 1986; pp 370-373.


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Table 4. Hydrogen Permeability of Various Materials material permeated

hydrogen isotope

K0 (mol/(m‚s‚Pa0.5) or mol/(m‚s‚Pa))

θK (K)

extrapolated K at 300 K, 1 Pa (mol/(m‚s))

measured temp range (K)

Fe stainless, 304S Ag Cu silica glass stainless, 309S M ) 15a glass Al M ) 30a glass Au Si β-SiC

H2 H2 H2 H2 H2 D2 H2 T2 H2 D2 H2 T2

4.1 E-8b 2.0 E-6 3.4 E-8 8.4 E-7 3.4 E-17‚T 1.5 E-7 6.1 E-17‚T 5.8 E-5 2.5 E-16‚T 3.1 E-6 1.4 E-5 1.8 E-10

4200 8660 7350 9320 3600 8500 6080 14800 8550 14800 27000 55600

3 E-14 6 E-19 8 E-19 3 E-20 6 E-20 7 E-20 3 E-23 2 E-26 3 E-26 1 E-27 1 E-44 6 E-91

375-850 370-570 730-980 470-700 300-1000 520-720 550-720 420-520 550-720 500-900 1240-1485 700-1570

a M is the percent of non-network formers (e.g., Na O, CaO, K O, and Al O ) which break up the network formers (e.g., SiO , B O , and 2 2 2 3 2 2 3 P2O5) that hold the glass together. b E-n ≡ × 10-n.

materials (having the largest negative exponent) appearing at the bottom of the table. Permeability is listed in mol/(m2‚s‚Pa0.5) for metals and mol/(m2‚s‚Pa) for glasses. The different pressure dependencies of these two classes of materials are due to the assumption that hydrogen will dissociate and diffuse through metals as atoms, whereas hydrogen diffuses through glasses as molecules. The table lists the hydrogen isotope that was used for measurements in each of the materials. A correction should be applied to estimate permeability for a desired isotope, since permeability increases as molecular weight to the -0.5 power. However, this correction is negligible compared to the error of extrapolating elevated temperature measurements to room temperature.7 From this table, one can deduce that the permeability of gold (or aluminum) is such that even a few contiguous monolayers could result in an acceptable hydrogen permeation barrier at room temperature, provided that such layers lacked imperfections (pinholes or cracks). Silicon or beta silicon carbide should be many orders of magnitude better. Copper or silver would be adequate if thin films are considered, instead of monolayers. Several rounds of American Society for Testing and Materials (ASTM) permeation barrier tests (described below) were used to test various metal layers deposited onto thin plastic substrates for use in fabricating lightweight bladders. These tests showed that defects such as pinholes and cracks can be more important than bulk properties when choosing a permeation barrier. In order to use thin films as permeation barriers in pressure vessels, other properties of the layer must be considered. The layers need to be formed or deposited into contoured liners with plumbing attachments (bosses), and they must be compatible with the range of temperatures, pressures, shock loads, strains, and chemicals that may be experienced from fabrication through the entire lifecycle of the pressure vessel. The liner may need additional properties to enable it to serve as a mandrel for composite overwrap. Pressure Vessels. The burst pressure (P) multiplied by internal volume (V) is a measure of the ultimate gas capacity of a pressure vessel. Dividing this product by tank weight (W) gives the performance factor (PV/W, usually measured in inches) for the pressure vessel. PV/W is a function of the properties of the materials used for the vessel’s fabrication. Lightweight pressure vessels with large PV/W will require considerable de-

Figure 4. PV/W for high cycle life pressure cylinders.17

velopment for URFC-powered vehicles, since hydrogen is not easily contained by composites or many other materials. Larger volume tanks, with higher burst pressures, and relatively small gas ports (bosses) will tend to have the highest ratio of high strength composite to relatively low strength liner materials (sufficient to limit hydrogen permeation), and so they achieve the highest performance factors. Aerospace tanks built with composite overwrapping on thin aluminum or titanium liners can be lightweight, but the highest performance factor designs are achievable only with relatively large capacity tanks, with low cycle life requirements. Metal liners will yield before their high-strength composite overwrap fails, causing cycle life to be limited by fatigue of the thin metal liner. For higher cycle life, thicker metal liners are required (at the expense of weight). Figure 4 shows how the use of laminated metallized bladder liners (to replace thick plastic or metal liners in graphite composite cylinders) can increase specific energy by ∼30% over comparable composite tanks designed for high cycle life. Bladder-lined composite pressure vessels can be an order-of-magnitude lighter weight compared to steel tanks for a similar capacity. These innovative tanks can be enabling to weight critical applications, such as HALE aircraft, pressurefed rocket propulsion, and other mobile or portable systems requiring tankage. Several thermal issues must be considered to design pressure vessels capable of fast filling with hydrogen.


Gas compression during pressure vessel filling results in gas heating which can result in underfilling and/or overpressurization. The temperature rise associated with filling a pressure vessel is related to the pressure ratio, fill rate, properties of the gas, thermal mass of the vessel and plumbing, and heat transfer coefficients. Large pressure ratios, rapid fill rates, hydrogen gas, and lightweight pressure vessels with poor heat transfer coefficient can result in large temperature rises. Neglecting heat transfer and the thermal mass of the vessel and plumbing, a study8 shows that for an infinite pressure ratio of an ideal gas, the temperature ratio for a fast fill is equal to the ratio of specific heats (which is 1.41 for hydrogen). By this criterion, a vessel with 300 K (27 °C ) 81 °F) initial gas temperature would achieve a final gas temperature of up to 420 K (147 °C ) 297 °F). Such worst case temperature rise would result in filling the tank to only 71.4% of rated capacity at the maximum operating pressure. The final gas temperature peaks can be worsened by high ambient temperature or gas lines that are already warm. Temperature rise (underfilling) for natural gas vessels is less than for hydrogen due to lower ratio of specific heats (∼1.32). Even so, by the criterion above, a fast fill from 300 K would yield a temperature of 396 K (253 °F). A typical temperature rise to ∼140 °F (333 K), from an initial temperature of 70 °F, for fast fills of compressed natural gas (CNG) into a vessel with an adiabatic inner wall has been reported.9 One of the main reasons for the discrepancy is due to the heat capacity of the pressure vessel, as will be discussed in a fast fill experiment described below. URFCs. URFCs should be capable of higher specific energy and less complexity (due to reduced parts count) compared with RFCs using separate (dedicated) fuel cell and electrolyzer stacks. However, there is an erroneous presumption that URFCs are not capable of high cycle life, due to the lack of well-publicized data showing tests lasting more than ∼25 cycles. Results from tests reported herein are an existence proof that URFC cells can be cycled repeatedly (>2000 cycles) without significant degradation (less than a few percent). These tests are also an existence proof that bifunctional catalysts can operate reversibly without significant degradation. Even though bifunctional catalysts will not be simultaneously optimized for both oxidation and reduction reactions, URFC systems will not necessarily be less efficient than RFC systems using dedicated stacks. Although it is true that bifunctional catalysts may have slightly reduced performance in either the oxidation or reduction reaction (depending on catalyst composition, which may be optimized for a given system application), this will result in a modest round-trip efficiency decrease per unit of active area for the URFC stack, and not necessarily an efficiency decrease for the entire URFC system. It should be realized that the efficiency of a dedicated fuel cell during the charge cycle is zero, and likewise, the efficiency of a dedicated electrolyzer during the discharge cycle is also zero. Therefore, it is more reasonable to compare the efficiency of RFC systems which use similar active areas. By this criter(8) Daney, D. E.; Edeskuty, F. J.; Daugherty, M. A.; Prenger, F. C.; Hill, D. D. Cryogenic Eng. Conf., July 17-21, 1995. (9) Kountz, K. J. Proc. 207th ACSsDivision of Fuel Chemistry, March 13-17, 1994.


Figure 5. URFC electrochemical cycles.

ion, the active area for the URFC stack is equivalent to the sum of the active areas for the dedicated FC and EC stacks combined. In such a comparison, a given power setting (either input or output) will result in less power per unit active area for the URFC by virtue of the larger active area utilized in each mode of operation. The operational current densities in the URFC will be below the corresponding current densities in the dedicated FC/EC stacks, resulting in an improved voltage efficiency for the URFC system. This efficiency improvement may even overwhelm the efficiency reduction caused by using compromise catalysts. The electrochemical reactions for FC and EC operations for hydrogen/oxygen based systems are depicted in Figure 5. The URFC can be designed for electrolysis with water fed from either the anode (oxygen side) or the cathode (hydrogen side). For cathode feed electrolysis, a single phase separator is used in the hydrogen/ water recirculation loop. Water must then diffuse through the cell to the oxygen side in order to be split. This necessary flow is decreased by an opposing flow of water caused by proton pumping, whereby each proton migrating through the membrane toward the cathode must be accompanied by at least four water molecules.10 This creates a situation where cell drying at the anode may result, especially at high current densities. Anode feed electrolysis provides excess water at the anode, avoiding the drying (at high current densities) that may be caused by proton pumping. For anode feed, phase separators are required in both the anode/water and the cathode/water recirculation loops, which add complexity and weight to the system. Other URFC chemistries are possible, such as hydrogen/halogen.11-14 These are of interest because they are capable of higher round-trip electrical efficiencies. Since they may be more than an order of magnitude heavier for comparable energy content, they are not of interest for mobile applications but are considered for stationary applications. Since halogens and their acids are corrosive and toxic, safety considerations may limit the use of hydrogen/halogen URFCs to specific niches where the improved efficiency is an overriding (10) Appleby, A. J.; Yeager, E. B. Energy 1986, 11, 137. (11) McElroy, J.; Patwa, G. G. Proc. 72nd AICHE Nov. 25-29 1979. (12) McElroy, J. Proceedings of the DOE Chemical Energy Storage and Hydrogen Energy Systems Contract Review; JPL Publication 781, November 16-17, 1977, 27-33. (13) Beaufrere, A.; Yeo, R. S.; Srinivasan, S.; McElroy, J.; Hart, G. 12th IECEC 1977, 959-963. (14) Nutall, L. J. “Development progress on solid polymer electrolyte fuel cells with various reactant sources”; General Electric, Electrochemical Energy Conversion Programs, Report, 1982.


Figure 6. Representative URFC polarization curves for various HBr concentrations.11

Figure 7. HBr URFC polarization curves demonstrated and projected.11

consideration. The H2O cycle, HCl cycle, and HBr cycle URFCs are shown schematically in Figure 5. Efficiency of hydrogen/halogen URFC systems is a function of acid concentration and current density, as illustrated by the representative polarization curves shown in Figure 6.11 This figure shows that unlike for water cycle URFCs, hydrogen/halogen stacks can be nearly 100% efficient, if operation is restricted to very low current densities, where FC and EC polarization curves approach one another. Figure 711 shows results from the late 1970s, where >90% electric-to-electric (ETE) voltage efficiency was achieved at low current density (108 mA/cm2). It was estimated that system-level parasitics would reduce the actual round-trip energy storage efficiency by ∼10% due to pumping power, current inefficiencies, and power conditioning inefficiencies. Electric utilities are interested in peak shaving energy storage systems to maximize utilization of existing base load electric generators and to postpone the installation of new generating equipment. Lead acid batteries are a preferred utility energy storage technology. Unlike the lead acid battery, RFCs uncouple power and energy ratings. This allows the RFC to accommodate weekly and even seasonal cycles. The H2/bromine URFC has been demonstrated in single cells for up to 4000 h with 80% round-trip energy storage efficiency, and showed

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stable cyclic performance in the early 1980s for a Boeing MX missile program.14 Scale-up of cells from 0.23 m2 (2.5 ft2) and minimization of corrosion currents have yet to be demonstrated. Less conventional stationary applications currently use lead acid batteries to store electric energy for remote off-grid applications. Primary electricity for these applications can come from solar, wind, or diesel generators. Up to ∼3 days of energy capacity are typically stored in the lead acid systems. H2/O2 or H2/air URFC systems will have longer life, and could have lower lifecycle cost, as well as allowing weekly and seasonal energy storage capabilities. The selection of H2/O2 versus H2/air URFCs depends on a number of considerations. URFC systems based on H2/O2 have higher performance per unit area of membrane, do not need oxidant compressors, can operate completely closed cycle with little maintenance, are capable of slightly higher efficiency, but require ∼50% more tankage to store oxygen. H2/O2 systems are generally preferred in cases where the compressed oxygen storage does not pose a significant safety hazard. Daytime shortages of electrical power (brown-outs or black-outs) are becoming more frequent in industrial countries. Total energy supplies are adequate on average, but daytime peaks cannot be sustained. A small H2/air URFC could provide the utility customer with energy during peak power periods. With sufficient numbers of individual homeowner units in place, this significant pressure on electric utilities would be reduced. Hydrogen/halogen URFC systems with an attractive mix of high efficiency and low capital cost for utility load leveling are being investigated enroute to a demonstration facility. URFC systems with lightweight pressure vessels have been designed for automobiles and are expected to be cost competitive with primary FC powered vehicles that operate on hydrogen/air with capacitors or batteries for power peaking and regenerative braking. URFCpowered vehicles can be safely and rapidly (250 °F ) 120 °C) suggest that bladder liners can be compatible with fast


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Figure 11. Inflated bladder with first composite wrap.

Figure 14. Wingspar load test photograph at EDO.

Figure 12. Inflated bladder with last composite wrap.

Figure 15. Wingspar test rig sketches. Figure 13. Fast fill experimental schematic.

fills and high-temperature curing cycles. Some of these materials are also compatible with cryogenic operation. A fast fill experiment at LLNL used a hydrogen source (deuterium) at 15 150 psi (104.5 MPa) to fill a 1.5 ft3 (0.0425 m3) stainless steel vessel (with maximum allowable working pressure of 6000 psi ) 41.4 MPa). The initial pressure for the steel vessel was atmospheric 15 psi (0.103 MPa) and it was filled to 3500 psi (24.1 MPa) in 5 min using a 2 stage cascade (Figure 13). After ∼20 h the pressure was 3226 psi and temperature was 293 K (20 °C ) 68 °F). The implied increase in absolute temperature (overpressurizing or underfilling) during the fill cycle was ∼8.5%, compared to the adiabatic limit of ∼41%. Structural Tanks. Representative sections of hybrid wingspar/tanks for a ∼200 ft (61 m) wingspan SRA were designed, fabricated, and load tested to failure in bending. The first series of experiments were performed on the apparatus shown in Figure 14, and shown functionally in the top rig sketched in Figure 15 (EDO Test Rig). Failure was due to subarea buckling in all cases. The first buckling failure was associated with the discontinuity conditions at the reinforcing plugs used to couple the test section to the load frame. In subsequent tests, this transition region was reinforced. Although the applied bending moment was correct, the equivalent tensile force on axis reduced the compression by ∼30%. Second and third generation test rigs (shown

as BYU Rig #1 and #2 in Figure 15) were constructed to correctly represent the wingspar loading conditions. The BYU Rig #1 failed in the support structure when a test article exceeded the 58 000 ft‚lb bending moment requirement by ∼5%. The BYU Rig #2 was reinforced to enable the structure to withstand >100 000 ft‚lb. Several design/fabricate/test iterations were required to exceed the safety factor of 2 requirement applied to the +5g flight moments (29 000 ft‚lb × 2 ) 58 000 ft‚lb or 79 000 N‚m), as presented in Table 3. The first successful composite prototype wingspar section weighs less than 55% of estimates for a similar section fabricated from high-strength aluminum capable of withstanding the same loads. Note that the +5g flight moment dominates the structural design and that this specification implies an order of magnitude higher moment than the spar designs being used on the current Pathfinder aircraft. URFCs. A photograph of part of the URFC test rig used at General Electric (GE) in 1972 is shown in Figure 16.15 The URFC cell tested in this rig was a prototype for use in zero gravity. In order to feed water to the cell for this design, wicking cloths were used and had not been optimized. Initially hydrophilic wicks became hydrophobic, so the sparse data recorded on this cell may be inconclusive. Figure 1715 shows the extent of (15) Chludzinski, P. J.; Danzig, I. F.; Fickett, A. P.; Craft, D. W. General Electric Company Technical Report AFAPL-TR-73-34, June 1973.



Figure 16. URFC cycle test at GE in 1972.15

Figure 17. URFC cycle test data from 1972.15 Figure 18. LLNL URFC test rig photograph.

the data that had been reported about this experiment. These data show that less than 40 mV degradation in both fuel cell and electrolyzer operation was observed when a single URFC cell was cycled ∼700 times at low current density (108 mA/cm2). A new URFC test facility has been built at LLNL. A primary fuel cell test rig with a single cell (46 cm2 active area) was modified for use as a URFC test rig. A photograph of the URFC test rig is shown in Figure 18 and is presented schematically in Figure 19. Hydrogen and oxygen are supplied by source bottles, or by recycle bottles that can be filled by electrolysis. A phase separator is used on the hydrogen side of the cell to separate hydrogen gas from a hydrogen/water mixture during cathode feed electrolysis. This rig was fabricated by enhancing the functionality of a primary fuel cell test rig. The URFC cell in this rig is located in the cabinet at the lower right of Figure 18, and is photographed separately in Figure 20. Experiments with this URFC test rig were limited within a parameter space of “safe” testing that could be performed while unattended. Exact duplication of the conditions in the 1972 study (108 mA/cm2 current density, room temperature, 0.45-2.17 MPa) was not possible because the LLNL apparatus was not certified to the same high pressure levels, and was not initially approved for gas capture during electrolysis. Detailed studies of degradation as a function of cycles demanded an accelerated test procedure that allowed measurements to be rapidly accumulated. Accumulating experience gained in debugging the apparatus verified that data could be logged periodically in a consistent, repeatable fashion with sufficient precision and accuracy to

Figure 19. LLNL URFC test rig schematic.

track small changes in performance as a function of cycle life (hundreds of microvolts to tens of millivolts). Operating temperatures and pressures were selected that could be safely and consistently maintained by the unattended apparatus during the protracted periods required to accumulate many cycles, representing conditions that would be of interest for several applications. The maximum allowable working pressure (MAWP) of the hydrogen side of the system is 80 psig (0.653 MPa), and the MAWP for the oxygen side is 160 psig (1.20 MPa). For safety reasons, the oxygen pressure was required to be slightly higher (consistently) than the hydrogen pressure. Since 0.45 MPa was nearly the minimum pressure in the 1972 data, and just below the


Mitlitsky et al.

Figure 21. URFC cycle test at LLNL for 2010 cycles.

Figure 20. LLNL URFC cell photograph.

pressure certification limit that bounded the safe (unattended) operational regime (in late 1996), 0.45 MPa was chosen for the oxygen side (oxygen plus water vapor). For safety purposes and unattended operational margin, 0.37 MPa was chosen for the hydrogen side (hydrogen plus water vapor). Various operational temperatures were considered between ∼24 °C (matching the 1972 GE data) and ∼90 °C (the maximum temperature allowed by some of the solenoids in the system). Unattended operation at this early stage made it unwise to stress the upper end of this range. An operating temperature of 82 °C was considered to be a useful temperature for some aircraft and automotive applications. High overall round-trip efficiency for closed systems is a strong function of diffusion losses, which increase exponentially with temperature. Therefore, relatively low operating temperatures, such as 49 °C (or less) may have merit. Cycling at 49 °C was specified anticipating efficiency measurements, as well as a desire to test temperature control. The H2 and O2 gas flows in the test rig are plumbed for flow through (rather than dead-ended) FC operation. High flow rates can cause some gas to flow past the membrane and electrode assembly (M&E) without reacting. Flow and pressure settings in the test rig were initially somewhat variable due to lack of feedback control. Data logging accuracy can be improved by measuring under conditions of high gas flow, in order to make measurements insensitive to slight flow variations. Initially, current density during cycling was investigated at 100 ASF (108 mA/cm2) to be consistent with 1972 GE data. However, electrolyzer current density for some RFC designs often optimizes at values higher than fuel cell current density. During some initial rapid cycle testing, it was determined that at least 200-240 ASF (215-258 mA/cm2) was required to quickly pressurize the system during electrolysis, to consistently maintain control of the positive pressure differential required on the oxygen during unattended operation, and to rapidly dry the hydrogen input side of the cell when switching from electrolysis to fuel cell operation. This short drying cycle may be required to remove the small amount of water that has been trapped in the “dead space” of the cell after water supply has been

turned off. For cells requiring very rapid switching times, this dead space should be minimized to reduce drying time. For the cell configuration being tested (which was a designed initially for dedicated fuel cell operation and converted to a URFC without considering this design parameter), it was determined that 2.0-2.5 min at 240 ASF (250 mA/cm2) were required for drying. Various cycle times were considered. Lunar rovers have cycle times of ∼28 days. HALE SRA have cycle times of ∼24 h. Low earth orbit satellites (and the 1972 GE experiment) have one cycle per ∼90 min. Automobiles using electrolytic regenerative braking may undergo charge cycles as short as 10-15 s. A 2.5 min electrolytic drying time was chosen and was set as half of the electrolysis (charge) cycle. A 5 min fuel cell (discharge) time and a 2.5 min cathode feed electrolysis time were chosen for a total cycle time of 10 min. Cycling of the URFC (cell LLNL01) was stopped periodically to acquire data for polarization curves. The data for these polarization curves were measured after allowing the system to stabilize for 30 min with high gas flow rates at 100 ASF (108 mA/cm2) fuel cell or 240 ASF (258 mA/cm2) electrolyzer. After 30 min, the system was quickly ramped to the highest current density to be recorded and allowed to stabilize for 2 min, and the voltage was recorded. The current density was then reduced by 20 ASF for fuel cell or 40 ASF for electrolyzer, followed by a 2 min stabilization period and data were again recorded. This was repeated until 0 ASF was recorded. The system was then rapidly ramped back to 100 ASF fuel cell or 240 ASF electrolyzer and another set of measurements was made to check drift. After fuel cell and electrolyzer polarization data were recorded, cycle testing was resumed. Cycle testing of a single-cell URFC was started in November 1996 at LLNL. The test results for a single cell cycled 2010 times are presented in Figure 21 at four different current densities for both fuel cell (FC) and electrolyzer (EC) modes (eight curves total). This experiment accurately measured the cell voltage under repeatable conditions to determine the extent of cell degradation. Zero degradation would be reflected in horizontal curves (no change in voltage as a function of cycle life). If there was significant degradation due to cycling, the FC curves would display larger negative slopes and the EC curves would display larger positive slopes. The plot in Figure 21 shows negligible degradation in both FC and EC performance, over a range of current densities. Six of the eight curves from Figure 21 are magnified in Figures 22-27. The plot limits for these expanded



Figure 22. URFC cycle test for FC at 43 mA/cm2.

Figure 26. URFC cycle test for EC at 172 mA/cm2.




Figure 28. URFC polarization curves after 750 cycles.


curves were chosen to show up to 50 mV of degradation and up to 30 mV of performance improvement compared to the cycle 1 measurement (80 mV full scale). A large fraction of remaining uncertainty about performance degradation is due to experimental error, which can be reduced with an upgraded apparatus. These figures show that performance has degraded by less than a few

percent during cycling. Figure 28 shows polarization curves measured for this cell after 750 cycles at two different temperatures for fuel cell and electrolyzer. Performance of this cell was limited by a large internal resistance (12 mohm/46 cm2 active area at 297 K), which dropped slightly to 11.4 mohm at 297 K by the end of the test. A new URFC membrane and electrode assembly (M&E) was installed and tested (cell 9734A). This cell had an internal resistance of 6.3 mohm/46 cm2 active area at 297 K, a value much closer to the expected value for a good cell. Internal resistance was measured to be 3.8 mohm/46 cm2 at 188 °F inlet temperature (360 K). A polarization curve for this cell is shown in Figure 29 at conditions which are very close to the highest temperatures and pressures allowable for the test rig in its present configuration. This figure proves that URFC cells can be cycled at >1 A/cm2 in both fuel cell and electrolyzer modes using Nafion 117 membrane. The polarization curves of Figure 28 and Figure 29 are plotted on the same graph in Figure 30 to show the


Mitlitsky et al.

Discussion

Figure 29. High-performance URFC polarization curves.

Figure 30. URFC polarization curve comparison.

performance improvement achieved. Cell 9734A has been operated as a fuel cell at current densities in excess of 1000 ASF (1.08 A/cm2, a factor of 3 higher than cell LLNL01). Cell 9734A was also operated to perform electrolysis at the same high power density, regenerating reactant gases for extended periods, despite the concern that cathode feed designs may be prone to membrane dry-out at high current density. These results are an important milestone and a significant improvement over what has previously been demonstrated. Both of the cells (9734A and LLNL01) tested used Nafion 117 membrane and Hamilton Standard’s E-5 catalyst, which is a proprietary mixture of platinum, platinum group metals, and their oxides. Catalyst loading was 4 mg/cm2 per electrode in both cells. The reduced internal resistance of cell 9734A and its corresponding performance improvement is attributed to the use of a new porous plate. The drying procedure required to cycle into FC operation has been accomplished in 25-30 s using a current density of 1.08 A/cm2. The drying procedure has also been accomplished in ∼15 s using a current density of 1.5 A/cm2. It is expected that redesign of the cell, catalyst, and drying procedure could reduce this drying time to a fraction of a second. Preliminary experiments also demonstrated that rapid cycling (with round-trip cycle times of less than 60 s) at current densities in excess of 1000 ASF is feasible. Two new URFC M&Es using Nafion 115 have been received by LLNL from Hamilton Standard. These cells have internal resistances of 5.7-5.8 mohm at room temperature and are expected to outperform the Nafion 117 cells, but have yet to be tested.

Permeation Barriers. ASTM hydrogen gas transmission tests were performed on laminates using gold, aluminum, and silver as the barrier layers. Although gold and aluminum have room temperature permeabilities that are 7-8 orders of magnitude lower than silver,7 measurements on laminates at room temperature showed that silver is a better choice as a barrier material. Permeation rates for silver laminates were consistent (within an order of magnitude) with the extrapolated room temperature permeability of silver, whereas permeation rates for gold and aluminum were >9 orders of magnitude worse than their respective extrapolated room temperature permeabilities. Hydrogen permeation in the gold and aluminum laminates was dominated by pinholes (or cracks for thicker aluminum films). A complex laminate with two thin aluminum layers in series had permeation that was 1-2 orders of magnitude lower than a single aluminum layer. Although this permeation rate is ∼50% higher than that of the silver laminate, it proves the utility of multiple film layers in reducing permeation dominated by pinholes. Such complex laminates may pose thermomechanical challenges to seam formation, since interlaminate bonds must be at least as strong as the seams in order to avoid failures due to delamination. The P-03 bladder material was developed for applications which had peak service temperatures of ∼140 °F (60 °C). One of the layers in the P-03 bladder has a peak temperature capability of ∼195 °F (90 °C). To better tolerate the heating that occurs during fast fills, and/or to incorporate some epoxy curing processes which may require temperatures to 250 °F (120 °C), a search for new bladder materials with higher temperature capability is underway. Bladder materials with upper service temperature >120 °C must be tested in preparation for forming into laminates. The exponential dependence of permeation on temperature requires a careful consideration of the highest temperatures expected during normal operation, including fast fills. Fully integrated bladder-lined pressure vessels have yet to be manufactured and tested in conformance with the standards required for their intended use. Hydrogen permeance is estimated to be less than 0.25%/year (∼8 g/yr out of ∼3.5 kg/aircraft) for a highaltitude aircraft point design using P-03 laminate for bladder liners.5 Scaling these room temperature results for automotive applications, which may have ∼20 times higher pressure (a square root dependence) and ∼10 times lower surface area (a linear dependence), suggests permeance comparable to high-altitude aircraft. The “handling” test provided early warning of the risk posed by wrinkles that might be captured in the liner during either the fiber winding or curing process steps. Although severe handling was shown to increase permeation locally by a factor of approximately 5, such degradation is not catastrophic for many applications and can be avoided by appropriate quality control during fabrication. Pressure Vessels. Prototype pressure vessels with PV/W of 40 km (1.6 million inches) were fabricated.2 Performance factors were calculated for automotive applications, investigating the tankage mass advantages of higher aspect ratio pressure vessels using



Table 5. PV/W for Bladder-Lined Hydrogen Storage Cylinders

a

fiber type

end dome thickness (in.)

boss wt per vessel (lb)

performance factor P(burst)V/W (in.)

H2 wt/(H2 + tank) wt (%)

T1000G T1000G T1000G T1000G T700S T700S T700S T700S T1000G T1000G T1000G T1000G

1/8 (0.317 cm) 1/4 (0.635 cm) 3/8 (0.952 cm) 1/2 (1.27 cm) 1/8 (0.317 cm) 1/4 (0.635 cm) 3/8 (0.952 cm) 1/2 (1.27 cm) 1/16 (0.159 cm) 1/8 (0.317 cm) 3/16 (0.476 cm) 1/4 (0.635 cm)

3.04 (1.38 kg) 3.04 (1.38 kg) 3.04 (1.38 kg) 3.04 (1.38 kg) 3.04 (1.38 kg) 3.04 (1.38 kg) 3.04 (1.38 kg) 3.04 (1.38 kg) 2.75 (1.25 kg) 2.00 (0.909 kg) 1.25 (0.568 kg) 0.52 (0.236 kg)

1.93 E6 (49.0 km)a 1.85 E6 (47.0 km) 1.77 E6 (45.0 km) 1.69 E6 (42.9 km) 1.54 E6 (39.1 km) 1.48 E6 (37.6 km) 1.43 E6 (36.3 km) 1.38 E6 (35.1 km) 2.00 E6 (50.8 km) 2.00 E6 (50.8 km) 2.00 E6 (50.8 km) 2.00 E6 (50.8 km)

12.4 12.0 11.5 11.1 10.2 9.8 9.5 9.2 12.8 12.8 12.8 12.8

En ≡ × 10n.

various fiber types, end domes, and bosses (see Table 5). The calculations are presented to three significant figures to show small variations in different designs. Actual performance factors are expected to be within a few percent of these values but require testing for validation. The last column of this table shows the % hydrogen storage by weight that is calculated by considering hydrogen weight/(vessel weight + hydrogen weight). The calculations assume three vessels per vehicle, 13.6 lb (6.18 kg) hydrogen per vehicle, 5 ksi (34.5 MPa) maximum operating pressure, 300 K operating temperature, 12.0 in. (30.5 cm) i.d., 51.0 in. (130 cm) overall length excluding boss protrusions, 2.25 safety factor, 924 ksi ultimate tensile strength of T1000G, 711 ksi ultimate tensile strength of T700S, ultimate tensile strengths derated to 81% of these values (accounting for variability in fibers and manufacturing), 1.80 g/cm3 fiber density, 1.20 g/cm3 resin density, 58% volume fraction of fibers, and 1.5:1 oblate spheroid approximate end dome shape. The last four entries in Table 5 show possible combinations of lightweighting the end domes and bosses that would yield a calculated PV/W ) 2.00 million in. (50.8 km) under the abovementioned assumptions. These tanks, when filled with hydrogen to their maximum operating pressure, would have a specific energy of 4260 Wh/kg based on 100% efficient utilization of the contained hydrogen’s lower heating value. The implied increase in absolute temperature (overpressurizing or underfilling) during the fill experiment (Figure 13) was ∼8.5%, compared to the adiabatic limit of ∼41%.8 This discrepancy is likely due to the assumptions made in calculating the adiabatic limit, which neglects heat flow into the pressure vessel wall (and the environment) during the filling process. The thermal mass of the 312 lb (142 kg) stainless steel test vessel was estimated to be 31 BTU/°F (59 kJ/K), whereas the thermal mass of the ∼3 lb (1.4 kg) of deuterium was estimated to be 5.2 BTU/°F (9.9 kJ/K). Although heat transfer to the environment during the fill process was probably small, the heat transfer into the wall of the cylinder was fast enough to substantially reduce temperature rise. High-strength composite pressure vessels with thin walls will provide thermal masses that are comparable to that of the hydrogen gas they contain, instead of being roughly 6-fold larger than hydrogen as in the experimental case. Therefore, temperature rises during fast fills of these lightweight pressure vessels will be accentuated, but still less than the adiabatic limit.

A number of strategies may be employed to compensate for heating during fast filling of pressure vessels. One study suggests precooling hydrogen gas to ∼214 K, to avoid underfilling, overpressurization, or high final gas temperatures.8 Precooling may be advantageous if liquid hydrogen is available at the filling station but could represent a large parasitic energy loss otherwise. Another study considers the use of phase change plugs in order to reduce temperature rise during fast fills.16 The materials used as the plugs would introduce significant weight and volume penalties. Active cooling with microchannels can solve heating problems but may be expensive and would add complexity to filling stations. Overpressurizing the vessel can be considered if tank materials are compatible with the worst case possible temperature rise. Safety factor would then be reduced for the time it takes to cool back to ambient temperature, which can be several hours for vessels with poor heat transfer. Heat transfer from the gas can be improved by using heat pipes, conducting foams, or other heat exchangers. The highest performance, lightest weight composite pressure vessels are currently available for large diameter, high pressure, low cycle life space applications. For applications requiring smaller diameter, lower pressure, and/or much higher cycle life, the performance factors will be significantly reduced. This performance factor reduction results from an increase in the ratio of parasitic masses (liners and bosses) in comparison to the mass of high-strength composite wall. In some cases, the liner weight dominates to the point where there is little merit in applying composite (as opposed to thickening the liner a little further and considering it a metallic pressure vessel). The use of thin bladder liners improves performance and enables composites to be competitive in relatively small diameter applications. For applications where weight and cost are paramount concerns, the weight decrease associated with the use of bladder liners can be traded for a weight gain accrued by using a lower strength fiber that is less expensive (e.g., using T700S or Panex 33 instead of T1000G). Structural Tanks. For some applications, the composite wall thickness required for pressure alone (or structure alone) can be less than the minimum gauge that is manufacturable and/or handleable. In such (16) Jasionowski, W.; Kountz, K. J.; Blazek, C. F. Gas Research Institute Report #GRI-92/0350, July 1992. (17) Arthur D. Little, Inc. “Multi-fuel reformers for fuel cells used in transportation: Assessment of hydrogen storage technologies”, Report Number DOE/CE/50343-1, March 1994.


cases, combining the functionality of pressure containment and structure into a single unit will enable a more efficient use of composite, with an associated weight savings. The HALE SRA design is an example of a pressure vessel with the desired shape designed to contain gas at the optimal pressure that would have specified a minimum gauge composite wall, causing difficulties with manufacturing and handling, were it not for the neighboring mass of structural composite. Combining the pressure containment requirements with the structural requirements of the wingspar results in a design which avoids the inefficiencies associated with minimum gauge fabrication and handling issues, while also providing additional structural integrity. The presence of internal pressure in a spar has the positive effect of increasing the buckling-critical moment. This improved structural performance cannot be considered an advantage in the design of the wingspar/tanks because of a global design requirement demanding that the structure remain intact even if there is a loss of internal pressure. In many cases, a tank designed for pressure loads may already be stiff enough to act as a structural member for the vehicle. In these cases there may be little or no additional composite required to act as a structural member, with the exception of designing nodes at the ends to couple the structural loads into the tank. For cases where structure requires slightly higher stiffness than is provided by the tank, alternate fibers with higher modulus and lower strength than T1000G (e.g., M40J) can be considered to help customize designs for minimum weight of the structural tank. The first successful composite prototype wingspar section weighs less than 55% of estimates for a similar section fabricated from high-strength aluminum capable of withstanding the same loads. Additional testing of wingspar/tanks under pressure and mixed loading conditions should be performed before integrating this design into a vehicle. URFCs. The plot in Figure 21 shows that negligible degradation occurred as a result of operating a URFC for 2010 cycles. The extent of degradation (if any) is less than 30 mV for all current densities reported and is dominated by experimental error in many cases. Ignoring the experimental error, the electrolyzer operation actually improved by a few millivolts over the course of cycling at all current densities reported. An upper limit for degradation can be calculated from Figures 22-27, by evaluating the difference between the most pessimistic extreme of the error bar on the final measurement (cycle 2010) and the most optimistic extreme of the error bars on the first measurement. Using this worst-case calculation, the upper limits for degradation were

Regenerative Fuel Cell Systems - ACS Publications - American

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