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Fast starting fuel processor for automotive fuel cell systems
International Journal of Hydrogen Energy 30 (2005) 953 – 962 www.elsevier.com/locate/ijhydene
Fast starting fuel processor for automotive fuel cell systems S.G. Goebel∗ , D.P. Miller, W.H. Pettit, M.D. Cartwright General Motors Corporation, Fuel Cell Activities, 10 Carriage St., Honeoye Falls, NY 14472, USA Received 2 September 2004; received in revised form 18 January 2005; accepted 20 January 2005 Available online 14 March 2005
Abstract A fuel processor was constructed which incorporated two burners with direct steam generation by water injection into the burner exhaust. These burners with direct water vaporization enabled rapid fuel processor start-up for automotive fuel cell systems. The fuel processor consisted of a conventional chain of reactors: auto-thermal reformer (ATR), water gas shift (WGS) reactor and preferential oxidation (PrOx) reactor. The criticality of steam to the fuel reforming process was illustrated. By utilizing direct vaporization of water, and hydrogen for catalyst light-off, excellent start performance was obtained with a start time of 20 s to 30% power and 140 s to full power. 䉷 2005 International Association for Hydrogen Energy. Published by Elsevier Ltd. All rights reserved. Keywords: Start-up; Fuel reformer; Hydrogen; Fuel cell
1. Introduction This work was part of an effort to develop on-board fuel processing for automotive fuel cell systems. For efficient, low-temperature fuel cells, hydrogen fuel or hydrogen rich reformate is required. While hydrogen storage on-board would provide operational simplicity, storing a sufficient quantity of hydrogen within the constraints of an automobile is problematic. The energy storage density and handling of a liquid fuel are much preferred. However, to meet customary automotive requirements, it is imperative that vehicles have a short start time. This work was specifically directed towards reducing the start-up time of a fuel processor for use with automotive fuel cell systems. The type of fuel processor considered was based on an auto-thermal reformer (ATR) for breakdown of the hydrocarbon fuel, water gas shift (WGS) reactor to increase the hydrogen content and a preferential oxidation (PrOx) ∗ Corresponding author. Tel.: +1 585 624 6730; fax: +1 585 624 6680. E-mail address: [email protected] (S.G. Goebel).
reactor to react CO that would poison low-temperature fuel cells. These reactors would typically be catalysts on monolith supports. The fuel processor would also include heat exchangers to reduce temperatures between reactors as well as to preheat reactants or to generate steam for the process.
2. Background The required amount of heat to achieve normal, fullpower operation can be expressed as the thermal energy of the components (m Cp Tr), where m is the mass of the component, Cp is the thermal heat capacity, and Tr is the temperature rise from ambient to operational temperature. This heating is typically obtained from the reactant gas flow (mf Cp Tf), where mf is the reactant mass flow, Cp is the thermal heat capacity, and Tf is the temperature change from reaction temperature to reactor exit. The reactors and heat exchangers are preferable designed with high surface area and low thermal mass. This creates relatively steep temperature gradients during heating. That is gases from the hot section of the reactor cannot travel very far into cooler sections without
0360-3199/$30.00 䉷 2005 International Association for Hydrogen Energy. Published by Elsevier Ltd. All rights reserved. doi:10.1016/j.ijhydene.2005.01.003
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Nomenclature Auto-thermal reformer atomic oxygen to carbon ratio preferential oxidation reactor molar steam to atomic carbon ratio water gas shift reactor
30 efficiency (%)
ATR o/c PrOx s/c WGS
P = 2.5 atm
ideal conditions 700 C CH4 equil 300 C WGS equil 1.5 O2/CO PrOx
35
conservative conditions 650 C CH4 equil 400 C WGS equil 3 O2/CO PrOx
25 20
9% H2
15 10
transferring their heat to the reactor. Therefore, the ratio of these terms, m Cp Tr/(mf Cp Tf), provides the heating time. To reduce the heating time, it is desirable to minimize the thermal mass of the components, increase the mass flow of reactants, and increase the temperature of the reaction. Decreasing component thermal mass also reduces the start-up energy, but the components need to be sufficiently sized to handle full-power operation and provide adequate longevity. The reactant mass flow would typically be limited by the flows available at full power. Increasing reaction temperature above the operating temperature is not effective. Excessively high temperatures could damage the components. Further it is necessary to achieve the operating temperature throughout the component. Although the higher reaction temperatures would more quickly provide the total energy requirement, the front portion of the reactor would be above the desired temperature and the downstream portion of the reactor would be cold. To get the entire reactor above the operating temperature, there is little benefit in operating the reaction above the normal operating temperature. The start time for the entire processor can be reduced by heating in sections. For example, the available flow could be used throughout the fuel processor with the first section operated fuel rich with stages of air addition downstream as disclosed by Robb and Pettit [1]. A fuel lean approach with stages of fuel addition is also possible, but it may be undesirable to switch between oxidative and reducing environments between start-up and normal operation, and the fuel would typically only be suitable for the ATR reactor. Therefore, a burner would need to be added for each additional stage as described by Goebel et al. [2]. During the start-up process, steam is normally not available. Steam would typically be generated by vaporizers between stages of PrOx reactors or downstream of the combustor. Until these reactors are heated and the exhaust from these reactors had heated the downstream vaporizers, steam cannot be generated by the conventional means. Therefore, the fuel processor heating process would typically have to be done without steam. Until stack grade reformate (meaning CO levels below about 100 ppm—and even at this level, fuel cell anode air bleed would still be required) is being generated, the overall fuel processor, including a combustor, must operate overall lean so as to not exhaust fuel. It is most desirable to operate close to stoichiometric conditions to obtain the maximum rate of energy input from the available air flow. Operation
5
2% H2
0 1
1.2
1.4
1.6
1.8 2 2.2 ATR o/c
2.4
2.6
2.8
3
Fig. 1. Efficiency without water from equilibrium.
near stoichiometric conditions is limited due to temperature considerations. The overall heat release is best accomplished by reaction in stages. Considering an overall stoichiometric temperature rise for hydrocarbon fuel and air of about 2000 ◦ C, and typically reaction temperatures of 700, 300 and 200 ◦ C for the ATR, WGS and PrOx, respectively, clearly several stages of reaction would be required. Further, when heat propagates to a downstream stage this will limit reaction in that stage. So the sections would ideally be sized to have the same heating time. With too few stages or mismatched section sizes, most of the air flow would go to the combustor to operate below its temperature limit leaving very little heating flow for the conversion portion of the fuel processor. The level of air flow used to start a fuel processor also needs to be considered. While it would be beneficial to use as much air flow as possible, in applications, approximately full power air flow at normal operating pressures would normally be the most available. Another issue with fuel rich start operation is either excessive temperature or carbon formation in the ATR. To avoid carbon formation, a molar air oxygen to fuel carbon ratio (o/c) greater than 1 would be required. However, at this o/c, the reaction temperature rise is about 900 ◦ C and an additional 400 ◦ C of preheat or more would be required to initiate the reaction. This temperature would lead to ATR catalyst sintering, and is well above the normal operating temperature of the downstream heat exchanger. It would be desirable to obtain operation without water due to freeze considerations for starting under sub-zero ambient conditions. However, equilibrium shows that very low efficiency would be obtained, and conservative operating conditions would push the efficiency even lower. This is illustrated in Fig. 1 for octene fuel. Considering an ATR equilibrium temperature (to establish a methane level), a WGS equilibrium temperature (to establish a CO level), and a PrOx O2 /CO ratio (to determine hydrogen consumption during the CO clean-up), the overall efficiency (heating value of output hydrogen to input fuel) was calculated. Ideal and conservative equilibrium
S.G. Goebel et al. / International Journal of Hydrogen Energy 30 (2005) 953 – 962
temperatures and PrOx O2 /CO ratios are given on the chart as a function of ATR o/c, with the result that the peak fuel processor efficiency would be 23–5% with a hydrogen concentration of 9% and 2%, respectively. Note that the conservative equilibrium temperatures also reflect that true equilibrium compositions are not achieved within the size of the reactor particularly as it is desirable to limit the size of the reactors to reduce the required start-up energy. Achieving ideal performance is also limited due to temperature and mixture non-uniformities within the reactors. Further, due to the low efficiency without water addition, sustained operation without water addition would lead to over temperature without extensive heat removal. Additional heat exchangers would exacerbate the problem of start-up energy. Also, for high overall ATR o/c levels, the air would need to be introduced in stages with intermediate heat removal to avoid excessive reaction temperatures, and the first stage would be prone to soot formation at the lower o/c levels of the first stage. An option for non-water operation is possible by using exhaust gas re-circulation as disclosed by Goebel and Pettit [3]. The dilution by the exhaust gas prevents excessive temperature for the high o/c operation and also provides product water from the fuel cell reaction. Most approaches to fuel processor start-up have been concerned with heating the various components to their operational temperature. However, it is possible to achieve operation before fully heating all of the components. In particular, it is possible to achieve operation by heating a portion of the reactors, but it is also required that steam be generated. Steam is normally generated by heat exchangers in the PrOx reactor or in the combustor exhaust. However, heat exchanger devices are typically more massive than the monolithic supported catalytic reactors for the same surface area due to the double wall construction to isolate the second fluid and the metal material typically used for heat exchangers. Further, these heat exchangers are typically placed downstream of the reactor, so their heating is delayed until the upstream reactor is nearly fully heated. So to achieve rapid start, an alternate source of steam is required during the initial phase of operation until the vaporizers have been heated. Then, for normal operation, it is more efficient to integrate heat from the process for reactant pre-heat or steam generation. The main difference of the fuel processor presented here from other fuel processors is the use of burners with downstream water spray to generate the steam needed for fuel processing to allow rapid fuel processor start-up. Additional options for this approach have been described by Goebel [4]. A risk of this approach could be damage to the catalysts due to water condensation; however, for any fast starting approach, water condensation must be tolerated. While the steam is necessary for fuel processing, it has the additional benefit of increasing the mass flow to heat the reactors. Steam is particularly effective, as it has a higher specific heat than air.
955
3. Experimental A fuel processor was constructed which incorporated two burners with direct steam generation. A schematic diagram is shown in Fig. 2 that has the same orientation as the photograph of the fuel processor in Fig. 3 with the first and second burners in the lower right and upper left corners, respectively. Flanges were used throughout the fuel processor so that any of the catalysts could be replaced. Also, the fuel processor was insulated, but the insulation was removed in the photograph. The layout of the burners is shown in Fig. 4. The fuel processor consisted of the conventional chain of reactors: ATR, WGS and PrOx, that were constructed from catalyst coated monoliths. A brief description of the components proceeding along the flow path follows. The first of the two burners was directed into the fuel processor inlet. The inlet used a solid-cone, direct-injection fuel injector that sprayed fuel co-axially with an air and steam flow. After a thin section of foam to isolate the ATR from the inlet, the first portion of the ATR consisted of conical shaped, catalyzed foam to diffuse the flow from the smaller inlet diameter followed by straight section of catalyzed monolith. A tube-in-shell heat exchanger was used to pre-heat the air and steam going to the inlet in counter-flow to the ATR outlet to cool the reformate before the first WGS reactor (WGS 1). A turn was located between the WGS stages, so a diffuser was required to redistribute the flow into the downstream WGS stage (WGS 2). A water injector was placed in the turn between the stages to provide a temperature reduction. The second burner was also tied into this location as the space provided some mixing between the burner exhaust and the flow from upstream. This second burner was used to directly supply steam to the second WGS and enable the fastest possible start time. Both WGS reactors were based on Platinum group metals rather than CuZn for improved temperature tolerance (especially for the first stage) and oxygen exposure tolerance to allow lean start methods. A two-stage, adiabatic PrOx reactor was used. Each PrOx stage included an array of tubes with holes for distributing the PrOx air followed by a vaporizer to cool the reformate and generate steam from the water supply, and a section of PrOx catalyst. A small heat exchanger in the reformate exit stream from the fuel processor pre-heated the water. This heat exchanger would also cool the reformate which would be needed before a membrane based fuel cell. The steam generated by the PrOx vaporizers was directed through a pressure regulator to the tube-in-shell heat exchanger mentioned previously along with the ATR air. Several drains were placed at key locations throughout the fuel processor to allow condensed water to be periodically drained during the start-up process by actuating the solenoid drain valves. Note that this fuel processor did not use a combustor that would normally burn the anode exhaust from a fuel cell. From a heat integration perspective, the combustor heat input is not required, but if available, could be used to generate additional steam.
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air
water
burn 2
fuel
air
WGS2
water
WGS 1
vap
HX
air
PrOx
ATR
vap
inlet
PrOx
fuel
HX
reformate
water
H2 air
water
burn 1
fuel
air
Fig. 2. Schematic of fuel processor.
Fig. 3. Photograph of fuel processor.
Each burner, as shown in Fig. 4, had an air line, a fuel nozzle, a spark plug for initiating the flame, and a water injector. The air was introduced through a swirler to stabilize the flame. A pressure atomizer was used to spray the fuel, and an injector solenoid was used to regulate the fuel flow. A similar arrangement was used to spray the water into the downstream section of the burner assembly. A conical section separated the flame zone from the water spray. This convergence ensured that the recirculation zone established by the swirler was closed and isolated the water spray from the flame. The cooled exhaust then flowed back over the burner wall to cool it from the flame before exiting. The burners, including the water vaporization sections, were 92 mm in diameter and 235 mm long. They were operated up to 0.43, 10 and 5.2 g/s of fuel, air and water, respectively. This represents a heat input rate of 18.6 kW.
To obtain the fastest possible start, hydrogen was used to ensure catalyst light-off. This light-off hydrogen was introduced into the line feeding the fuel processor inlet. An effort during the course of the testing was to use a minimal amount of hydrogen. Electrically heated catalyst sections could also be used to achieve catalyst light-off. A summary of the full power design conditions is provided in Table 1. The mass of the components is listed in Table 2. The diameter of the reactors was 144 mm except for the ATR where an 80 mm diameter monolith was initially used. In production, the flanges would not be required, so the total mass would be reduced to 26.7 kg. The reactor shells (the metal case and fibrous pad around the reactor monoliths) were also listed separately as this mass may not need to be fully heated to start the fuel processor. The external insulation mass was not included. It should be noted
to fuel processor
S.G. Goebel et al. / International Journal of Hydrogen Energy 30 (2005) 953 – 962
957
er
irl
sw wa sp ter ray
air supply
water injector
flame zone
fuel injector
water drain
spa plu rk g
Fig. 4. Schematic of start-up burner.
Table 1 Full power design conditions
Table 2 Component masses
Flows (g/s) Fuel ATR air PrOx air Water (PrOx) Water (WGS - inj)
2.0 8.2 1.4 4.6 1.0
Component
Ratios o/c (ATR) s/c (ATR) s/c (WGS 2) CO (mole%, WGS1in) CO (mole%, WGS2in) CO (mole%, WGS2out) O2/CO (PrOx)
0.83 1.8 2.2 9.3 3.7 0.6 2.0
Mass (kg)
Reactors Heat exchangers Inlet and mixing sections Reactor shells Manifolds and plumbing Burners Flanges Total
34.1
95 eff eff (CH4 cor) eff (CH4&NMHC cor)
90
68 78
that the total platinum group metal loading was 44.5 g with over three-quarters of this being used in the WGS reactors, so conversion to a CuZn WGS catalyst at least for the larger second stage could significantly reduce the precious metal content.
4. Results and discussion The fuel processor was initially evaluated under steady state conditions, and the results are shown in Fig. 5. The power and efficiency were calculated as outlined below. The power was based on the lower heating value of the hydro-
85 Efficiency (%)
Power, efficiency H2 (kW) Efficiency (%)
5.3 9.6 2.0 3.6 2.6 3.6 7.4
80 75 70 65 60 55 50 0
10
20
30 40 50 Power (kW H2)
60
70
Fig. 5. Efficiency over operating range.
gen from the fuel processor and the efficiency was normalized by the lower heating value of the input fuel. The fuel type used was a mixture of pure components to represent
S.G. Goebel et al. / International Journal of Hydrogen Energy 30 (2005) 953 – 962
1000
1.00 PrOx CO T ATR T WGS1 T WGS2 T PrOx1 T PrOx2 WGS2 CO
900
Temp (C), PrOx (ppm)
800 700 T ATR
600
0.90 0.80 0.70 0.60
WGS2 CO
500
0.50
400
0.40 T WGS1
T WGS2
300
WGS2 CO (%)
958
0.30 T PrOx1
200
0.20 T PrOx2
100
0.10 PrOx CO
0 0
20
40
60
80 100 time (s)
120
140
160
0.00 180
Fig. 6. Start-up temperatures and CO for test to half power.
70
power (kW H2)
60 50
h = (2 − (o/c) − 4a/x − 3a /x − 3b/x)x + y/2 (moles H2 per mole fuel assuming NMHC on C1 basis as Cx H2x with an O2/CO ratio of 1.5 for additional CO clean-up), (2)
Start Energy to Half Power hydrogen input = 7.5 g (1.8 MJ gasoline equiv. at 50% conv. and storage) gasoline input = 136 g (5.9 MJ) hydrogen output = 1.7 MJ (2.3 MJ gasoline equiv. at 75% reforming) start energy = 5.4 MJ (124 g or 0.7 cup) thermal mass (non-flanged) = 5.2 MJ WGS2 air on (to keep warm)
40 run fuel on (over-shoot) hydrogen off (when ATR hot enough for light-off)
30
burner 1 off (when steam from vaporizers)
20 10
burner 2 off (when steam from burner 1 to WGS2)
hydrogen and burner heating
0 0
20
40
60
80 100 time (s)
120
140
160
180
Fig. 7. Start-up power and transitions for test to half power.
a/x = [CH4 ]/([CO2 ] + [CH4] + [NMHC] + [CO]) (from GC measurements),
(3)
a /x = [NHMC]/([CO2 ] + [CH4] + [NMHC] + [CO]) (for NMHC on C1 basis from GC measurements), (4) b/x = [CO]/([CO2] + [CH4] + [NMHC] + [CO]) (from CO gas analyzer and GC measurements), (5)
the range of hydrocarbons found in gasoline. The fuel average composition was Cx Hy where x = 7.494 and y = 14.53 for a molecular weight of 104.5, and had a lower heating value of 4536 kJ/g mole. Metered flows were used for these calculations and corrections were made for any methane, non-methane hydrocarbon (NMHC), or carbon monoxide slip based on gas chromatograph (GC) and CO gas analyzer measurements (the a, a and b terms, respectively). For this fuel processor, as a combustor was not used, the anode stoichiometry did not need to be considered nor a compensation for loss in efficiency as with other fuel processors. Cx Hy + (o/c)x/2 O2 + (o/c)x/23.76 N2 + (s/c)xH2 O → (x − a − a − b)CO2 + h H2 + w H2 O + 0 O2 + a CH4 + a NMHC + b CO + (o/c)x/2 ∗ 3.76 N2 (chemical equation), (1)
efficiency(%) = h LHVH2 (kJ/g mole)/ LHV fuel(kJ/g mole) ∗ 100
(6)
for x = 7.494, y = 14.53, LHV H2 = 242 kJ/g mole, LHV fuel = 4536 kJ/g mole, efficiency(%) = 118.7 − 40 (o/c) (for a, a and b = 0), (7) power (kW H2 ) = efficiency(%)/100 ∗ fuel(g/s) ∗ LHVfuel/MWfuel,
(8)
o/c = air(g/s)/fuel(g/s) ∗ (2 ∗ 0.21/MWair) /(x/MWfuel),
(9)
s/c = water(g/s)/fuel(g/s) ∗ (1/MWwater) /(x/MWfuel)
(10)
S.G. Goebel et al. / International Journal of Hydrogen Energy 30 (2005) 953 – 962
1000
1.00 PrOx CO T ATR T WGS1 T WGS2 T PrOx1 T PrOx2 WGS2 CO
800 700
0.90 0.80 0.70
600
0.60
500
0.50 WGS2 CO
400
0.40
T WGS1 T WGS2
300
WGS2 CO (%)
T ATR
900 Temp (C), PrOx CO (ppm)
959
0.30
T PrOx2
200
0.20 T PrOx1
100
0.10 PrOx CO
0 0
20
40
60
80 100 time (s)
120
140
160
0.00 180
Fig. 8. Start-up temperatures and CO for test to full power.
for MWfuel = 104.5, MWair = 28.85, MWwater = 18, o/c = 0.203 air(g/s)/fuel(g/s),
(11)
s/c = 0.774 water(g/s)/fuel(g/s).
(12)
The efficiency data in Fig. 5 represents the entire range of fuel processor performance. The methane and NMHC slip from this fuel processor were excessively high leading to substantial losses in overall efficiency and power. If this problem were fixed, the efficiency would be close to the expected 80% value. It should be noted that some of the uncorrected efficiencies were high (almost 90%) due to low o/c operation to limit the temperature of hot regions in the ATR (which allowed for hydrocarbon slip in the cool regions). Much lower levels of methane and no NMHC slip were observed on previous tests with similar hardware. Additional evaluations and modifications were undertaken to correct this problem. Temperature profiles across the ATR did indicate non-uniformity under many conditions. At lower power, the inlet exhibited hollow cone injector characteristics with hotter temperatures at the center, and the spray may not have fully penetrated at high power leading to a more uniform temperature (or hotter temperatures at the wall in some cases). Additional 80◦ angle spray injectors were tried, but they continued to follow the previously observed patterns. A 60◦ angle spray injector was also tested and provided more uniform temperature profiles at low power but became un-runnable at higher powers due to high temperatures at the wall. So the 80◦ injector was used for the remainder of the runs. To facilitate mixing, additional catalyzed foam was added to the ATR and a larger monolith section was also used to improve fuel conversion. The replaced ATR catalyst significantly reduced the NMHC slip initially, but after a few tests, the high levels of NMHC slip returned.
Even under conditions where the temperature profile (inlet mixing) was uniform, high levels of hydrocarbon slip were still observed. From the GC speciation, it was noted that toluene (one of the parent molecules) contributed over 50% of the NMHC (on a C1 basis) while octane (the most abundant parent molecule) was not observed. This may indicate a preference of fuel type for fuel processing or poorer catalyst performance on certain types of fuel compounds. Another factor is the fuel pulsation from the injector (which would require storage capacity on the catalyst), however, at full power, the injector was held open. Other notable performance metrics were the PrOx performance and pressure drop. The PrOx reactor performed better than expected reducing CO levels to below 100 ppm with an O2/CO ratio of about 1.5 that was lower than the expected ratio of 2.0 given in Table 1. The pressure drop of the entire fuel processor was only 16 kPa at full power operation at a back-pressure of 50 kPa g. Figs. 6–9 show fuel processor starts using hydrogen for catalyst light-off. Shown are CO levels out of the PrOx and WGS 2 (labeled as PrOx CO and WGS 2 CO, respectively). Also shown are temperatures at the front of reactors. Hydrogen output versus time was calculated from the flows, and this information, as well as fuel input, was used to determine start energy. These charts also indicate the steps in the start sequence. The process was initiated by starting the air flows to Burner 1 and Burner 2 as well as the two PrOx reactors of 10, 5, 1.2 and 0.7 g/s, respectively. So the start-up air flow was somewhat greater than the normal operation air flow. Then the burner fuels were started and the burners were lit. After the burners were lit, the burner water sprays were turned on. The light-off hydrogen was started after burner operation was established. The primary function of the light-off hydrogen was to quickly get the
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S.G. Goebel et al. / International Journal of Hydrogen Energy 30 (2005) 953 – 962 70
50 40 30
run fuel on (over-shoot) hydrogen off (when ATR hot enough for light-off)
20 10 0 0
burner 1 off (when steam from vaporizers)
burner 2 off (when steam from burner 1 to WGS2)
hydrogen and burner heating
power (kW H2)
60
Start Energy to Full Power hydrogen input = 3.5 g (0.9 MJ gasoline equiv. at 50% conv. and full power storage) (when WGS1 gasoline input = 172 g (7.5 MJ) heated) hydrogen output = 3.3 MJ (4.4 MJ gasoline equiv. at 75% reforming) start energy = 3.9 MJ (90 g or 1/2 cup) thermal mass (non-flanged) = 5.2 MJ
20
40
60
80
100
120
140
160
180
time (s)
Fig. 9. Start-up power and transitions for test to full power.
front of the ATR to its light-off temperature, but the lightoff hydrogen was also important to initiate PrOx heating as these reactors were not directly heated by a burner. After the front of the ATR was above its light-off temperature (which was achieved in about 20 s after the start of the process for the test shown in Figs. 8 and 9), the hydrogen was shut off and the run fuel was started and the fuel processor begins to make reformate. It is desirable to use as little start-up hydrogen as possible to reduce the start-up energy requirements. The run fuel was set at a level commensurate with half power under normal operation, but due to the higher o/c operation during start-up, about 30% of full power hydrogen was generated. When the steam from Burner 1 was established through the fuel processor, Burner 2 and its water spray could be shut off. It is desirable to shut off the burners as soon as possible to reduce the start-up energy requirements. A reduced air flow was continued through Burner 2 to allow direct reformate-air reaction on WGS 2 to continue heating this reactor that would otherwise be cooled down as the upstream flow from WGS 1 was cooler at this point. When the PrOx vaporizers began to warm, water flow to these vaporizers was started. When vaporizers began to fully generate steam, Burner 1 and its water spray were shut off. Air for the ATR reaction was then supplied via the normal route through the heat exchanger rather than by the lean exhaust from Burner 1. At this point additional heating of WGS 1 was desired before transitioning to full power. Most of the starts showed a CO spike when the run fuel was started as the fuel flow typically overshot the desired value as the controls adjusted to the fuel pressure. CO spikes were often observed when Burner 1 was shut off as all the steam generation must be supported by the PrOx vaporizers. CO spikes were also observed upon the transition to full power. Fig. 6 shows a start with a longer hydrogen light-off period that had smaller CO spikes than other starts. To achieve low CO and a fast start to full power requires a good sequence of burner operation to sustain the needed steam (and not stopping burner steam generation before the PrOx vaporizers are produc-
ing adequate steam), appropriate air flows to sustain the reactor temperatures (without an over-temperature) and purging to remove condensed water. For the start shown in Fig. 8, the later portion had low CO, so in combination with the first portion of the start shown in Fig. 6, it appears possible to achieve a fast start with low CO throughout. From fuel cell testing, brief spikes over 1000 ppm of CO are tolerable as long as the average CO level can be handled. A drop in ATR temperature shows the need to purge excess water during the start sequence (this occurs when burner 1 is shut off as air then comes through the ATR HX which probably has condensed water in the bottom of the shell). These start-up sequences were performed using manually initiated steps. Additional controls development could improve the consistency of the start-up process. To put the start energy into perspective, it is compared to the energy used for driving. For example, consider a reformate based fuel cell vehicle that achieves 60 mpg during driving. It would require 30.7 MJ of fuel over a 15 mile drive cycle. For the fuel processor start energy, the fuel input, start-up hydrogen and additional air compressor work are considered. Fuel input for start-up was taken to be the amount above what was reformed. The amount of fuel reformed was based on the amount of hydrogen produced using a 75% conversion efficiency. The start-up hydrogen was included at twice the energy (to account for formation and storage efficiency) and some compression work (about 0.3 MJ on a fuel equivalent basis) was included to account for the additional air flow during start-up. On this basis, the fuel processor would require a fuel equivalent start energy of 4.2 MJ (for the start-up shown in Fig. 9). Compared to the fuel used for driving in this example, this is an additional 14% fuel. Including this additional start-up energy in the total drive cycle fuel consumption, the drive cycle average fuel economy would be reduced to 53 mpg. The thermal portion of the start energy of 3.5 MJ is significantly less than the thermal energy of 5.2 MJ to heat the entire fuel processor (excluding the flanges and burners) to operating conditions which indicates that the fuel processor can operate at full power conditions without being fully heated (probably the reactor shells are not fully heated). This also indicates that some of the normal operation waste heat (which typically appears as steam in the reformate output) can be used to heat the fuel processor during start-up. This also demonstrates that the energy used for this start-up process was effectively utilized. To store the start-up hydrogen, using a 5000 psig (341 MPa) tank, only 0.012 l per start would be required. Such a tank could be replaced at each fuel refill. Alternative start methods without using hydrogen were also investigated. One approach involves generating reformate by operating a burner fuel rich and requires that the catalysts light-off on this reformate. Reformate light-off capability was evaluated by establishing air flows and then Burner 1 was lit under fuel rich conditions (with water spray to control the burner temperature). Burner 1 was used to
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1000
1.00 PrOx CO T ATR T WGS1 T WGS2 T PrOx1 T PrOx2 WGS2 CO
800 700 600
0.90 0.80 0.70 0.60
T ATR
500
0.50
WGS2 CO
400
0.40
T WGS2
300
WGS2 CO (%)
900 Temp (C), PrOx CO (ppm)
961
0.30 T PrOx1
200
0.20
T PrOx2
100
T WGS1
0.10
PrOx CO
0
0.00 0
20
40
60
80
100 120 140 160 180 200 220 240 time (s)
Fig. 10. Start-up temperatures and CO for lean start.
70
power (kW H2)
60 50 40
Start Energy to Full Power hydrogen input = 0 g (non hydrogen start) gasoline input = 202 g (8.8 MJ) hydrogen output = 3.3 MJ (4.4 MJ gasoline equiv. at 75% reforming) start energy = 4.4 MJ (101 g or 0.6 cup) brief fuel increase thermal mass (non-flanged) = 5.6 MJ (to moderate ATR temperature) run fuel on (over-shoot) (when ATR hot enough for light-off)
30 20
full power (when WGS1 heated)
burner 2 and 1 off (when steam from vaporizers)
10 burner heating
0 0
20
40
60
80 100 120 140 160 180 200 220 240 time (s)
Fig. 11. Start-up power and transitions for lean start.
avoid any direct heating of the catalysts and to delay the arrival of the condensation wave that appeared as a temperature rise in WGS 2 about 15–20 s after the burner was lit. It was found that the WGS catalyst did and the PrOx catalyst did not light-off on fuel rich burner exhaust and air. Only after an inlet temperature rise did significant reactions begin to occur on the PrOx catalyst. Another concern with this approach was generation of soot by the fuel rich flame. Examination of the hardware after a few tests did show soot deposits on the reactors. Based on the PrOx catalyst not lighting-off and burner sooting, it was decided not to pursue fuel rich burner starts. Non-hydrogen start-ups were successfully done by operating both burners fuel lean as shown in Figs. 10 and 11. Here, the burners function to warm the reactors to their lightoff temperatures and also provide steam until the vaporizers are heated. The water spray also reduces the burner exhaust
temperature and increases the heating mass flow. The process was initiated by starting the air flows to Burner 1 and Burner 2 of 10 and 5 g/s, respectively. Then the burner fuels were started and the burners were lit. After the burners were lit, the burner water sprays were turned on. After the front of the ATR was above its light-off temperature (which was achieved about 50 s after the start of the process for the test shown in Figs. 10 and 11), the run fuel and PrOx air flows were started and the fuel processor begins to make reformate. The run fuel was set at a level somewhat below half power under normal operation to operate at a slightly higher o/c to assist heating of the ATR. When the PrOx vaporizers began to warm, water flow to these vaporizers was started. When the vaporizers began to generate steam, Burners 1 and 2 were shut off. The CO spike observed at this point may be due to the steam generation of the PrOx vaporizers being at a rate lower than that produced by the burners with water spray. Air for the ATR reaction was then supplied via the normal route through the heat exchanger rather than by the lean exhaust from Burner 1. A reduced air flow was continued through Burner 2 to allow direct reformate-air reaction on WGS 2 to continue heating this reactor that would otherwise be cooled down as the upstream flow from WGS 1 was cooler at this point. Additional heating of WGS 1 was desired before transitioning to full power. WGS 1 was slow to heat, so it may have been beneficial to extend the duration of Burner 1 operation.
5. Conclusions The criticality of steam to the fuel reforming process was illustrated. By utilizing direct vaporization of water in burner
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exhaust, and hydrogen for catalyst light-off, excellent start performance was obtained with a start time of 20 s to 30% power and 140 s to full power. The requirement to have water for a fuel processor start (and operation) clearly creates problems for automotive applications that must tolerate subfreezing conditions. Start methods not using light-off hydrogen were also tested. Issues with PrOx light-off and burner sooting were found with the fuel rich start method. A fuel lean start method to full power was demonstrated with 50 s to start and 190 s to full power. Again the use of water spray was beneficial to increase the heating mass flow and necessary to provide steam until the vaporizers were heated.
Acknowledgements A much larger team was working on fuel processing at General Motors, and they provided the catalysts, heat exchangers and build support for this fuel processor. The support of General Motors and its management team for the opportunity to demonstrate the capabilities of fuel processing for automotive fuel cell systems was also greatly appreciated.
References [1] Robb GM, Pettit WH. Method for quick start-up of a fuel processing system using controlled staged oxidation. US Patent Appl. US2003/0170510; 2003. [2] Goebel SG, Pettit WH, Sennoun MEH, Miller DP. Staged lean combustion for rapid start of a fuel processor. US Patent Appl. US2003/0093949; 2003. [3] Goebel SG, Pettit WH. Fuel processing system having gas recirculation for transient operations. US Patent Appl. US2003/0136880; 2003. [4] Goebel SG, Direct water vaporization for fuel processor startup and transients. US Patent Appl. 2003/0154654; 2003.
Fast starting fuel processor for automotive fuel cell systems S.G. Goebel∗ , D.P. Miller, W.H. Pettit, M.D. Cartwright General Motors Corporation, Fuel Cell Activities, 10 Carriage St., Honeoye Falls, NY 14472, USA Received 2 September 2004; received in revised form 18 January 2005; accepted 20 January 2005 Available online 14 March 2005
Abstract A fuel processor was constructed which incorporated two burners with direct steam generation by water injection into the burner exhaust. These burners with direct water vaporization enabled rapid fuel processor start-up for automotive fuel cell systems. The fuel processor consisted of a conventional chain of reactors: auto-thermal reformer (ATR), water gas shift (WGS) reactor and preferential oxidation (PrOx) reactor. The criticality of steam to the fuel reforming process was illustrated. By utilizing direct vaporization of water, and hydrogen for catalyst light-off, excellent start performance was obtained with a start time of 20 s to 30% power and 140 s to full power. 䉷 2005 International Association for Hydrogen Energy. Published by Elsevier Ltd. All rights reserved. Keywords: Start-up; Fuel reformer; Hydrogen; Fuel cell
1. Introduction This work was part of an effort to develop on-board fuel processing for automotive fuel cell systems. For efficient, low-temperature fuel cells, hydrogen fuel or hydrogen rich reformate is required. While hydrogen storage on-board would provide operational simplicity, storing a sufficient quantity of hydrogen within the constraints of an automobile is problematic. The energy storage density and handling of a liquid fuel are much preferred. However, to meet customary automotive requirements, it is imperative that vehicles have a short start time. This work was specifically directed towards reducing the start-up time of a fuel processor for use with automotive fuel cell systems. The type of fuel processor considered was based on an auto-thermal reformer (ATR) for breakdown of the hydrocarbon fuel, water gas shift (WGS) reactor to increase the hydrogen content and a preferential oxidation (PrOx) ∗ Corresponding author. Tel.: +1 585 624 6730; fax: +1 585 624 6680. E-mail address: [email protected] (S.G. Goebel).
reactor to react CO that would poison low-temperature fuel cells. These reactors would typically be catalysts on monolith supports. The fuel processor would also include heat exchangers to reduce temperatures between reactors as well as to preheat reactants or to generate steam for the process.
2. Background The required amount of heat to achieve normal, fullpower operation can be expressed as the thermal energy of the components (m Cp Tr), where m is the mass of the component, Cp is the thermal heat capacity, and Tr is the temperature rise from ambient to operational temperature. This heating is typically obtained from the reactant gas flow (mf Cp Tf), where mf is the reactant mass flow, Cp is the thermal heat capacity, and Tf is the temperature change from reaction temperature to reactor exit. The reactors and heat exchangers are preferable designed with high surface area and low thermal mass. This creates relatively steep temperature gradients during heating. That is gases from the hot section of the reactor cannot travel very far into cooler sections without
0360-3199/$30.00 䉷 2005 International Association for Hydrogen Energy. Published by Elsevier Ltd. All rights reserved. doi:10.1016/j.ijhydene.2005.01.003
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Nomenclature Auto-thermal reformer atomic oxygen to carbon ratio preferential oxidation reactor molar steam to atomic carbon ratio water gas shift reactor
30 efficiency (%)
ATR o/c PrOx s/c WGS
P = 2.5 atm
ideal conditions 700 C CH4 equil 300 C WGS equil 1.5 O2/CO PrOx
35
conservative conditions 650 C CH4 equil 400 C WGS equil 3 O2/CO PrOx
25 20
9% H2
15 10
transferring their heat to the reactor. Therefore, the ratio of these terms, m Cp Tr/(mf Cp Tf), provides the heating time. To reduce the heating time, it is desirable to minimize the thermal mass of the components, increase the mass flow of reactants, and increase the temperature of the reaction. Decreasing component thermal mass also reduces the start-up energy, but the components need to be sufficiently sized to handle full-power operation and provide adequate longevity. The reactant mass flow would typically be limited by the flows available at full power. Increasing reaction temperature above the operating temperature is not effective. Excessively high temperatures could damage the components. Further it is necessary to achieve the operating temperature throughout the component. Although the higher reaction temperatures would more quickly provide the total energy requirement, the front portion of the reactor would be above the desired temperature and the downstream portion of the reactor would be cold. To get the entire reactor above the operating temperature, there is little benefit in operating the reaction above the normal operating temperature. The start time for the entire processor can be reduced by heating in sections. For example, the available flow could be used throughout the fuel processor with the first section operated fuel rich with stages of air addition downstream as disclosed by Robb and Pettit [1]. A fuel lean approach with stages of fuel addition is also possible, but it may be undesirable to switch between oxidative and reducing environments between start-up and normal operation, and the fuel would typically only be suitable for the ATR reactor. Therefore, a burner would need to be added for each additional stage as described by Goebel et al. [2]. During the start-up process, steam is normally not available. Steam would typically be generated by vaporizers between stages of PrOx reactors or downstream of the combustor. Until these reactors are heated and the exhaust from these reactors had heated the downstream vaporizers, steam cannot be generated by the conventional means. Therefore, the fuel processor heating process would typically have to be done without steam. Until stack grade reformate (meaning CO levels below about 100 ppm—and even at this level, fuel cell anode air bleed would still be required) is being generated, the overall fuel processor, including a combustor, must operate overall lean so as to not exhaust fuel. It is most desirable to operate close to stoichiometric conditions to obtain the maximum rate of energy input from the available air flow. Operation
5
2% H2
0 1
1.2
1.4
1.6
1.8 2 2.2 ATR o/c
2.4
2.6
2.8
3
Fig. 1. Efficiency without water from equilibrium.
near stoichiometric conditions is limited due to temperature considerations. The overall heat release is best accomplished by reaction in stages. Considering an overall stoichiometric temperature rise for hydrocarbon fuel and air of about 2000 ◦ C, and typically reaction temperatures of 700, 300 and 200 ◦ C for the ATR, WGS and PrOx, respectively, clearly several stages of reaction would be required. Further, when heat propagates to a downstream stage this will limit reaction in that stage. So the sections would ideally be sized to have the same heating time. With too few stages or mismatched section sizes, most of the air flow would go to the combustor to operate below its temperature limit leaving very little heating flow for the conversion portion of the fuel processor. The level of air flow used to start a fuel processor also needs to be considered. While it would be beneficial to use as much air flow as possible, in applications, approximately full power air flow at normal operating pressures would normally be the most available. Another issue with fuel rich start operation is either excessive temperature or carbon formation in the ATR. To avoid carbon formation, a molar air oxygen to fuel carbon ratio (o/c) greater than 1 would be required. However, at this o/c, the reaction temperature rise is about 900 ◦ C and an additional 400 ◦ C of preheat or more would be required to initiate the reaction. This temperature would lead to ATR catalyst sintering, and is well above the normal operating temperature of the downstream heat exchanger. It would be desirable to obtain operation without water due to freeze considerations for starting under sub-zero ambient conditions. However, equilibrium shows that very low efficiency would be obtained, and conservative operating conditions would push the efficiency even lower. This is illustrated in Fig. 1 for octene fuel. Considering an ATR equilibrium temperature (to establish a methane level), a WGS equilibrium temperature (to establish a CO level), and a PrOx O2 /CO ratio (to determine hydrogen consumption during the CO clean-up), the overall efficiency (heating value of output hydrogen to input fuel) was calculated. Ideal and conservative equilibrium
S.G. Goebel et al. / International Journal of Hydrogen Energy 30 (2005) 953 – 962
temperatures and PrOx O2 /CO ratios are given on the chart as a function of ATR o/c, with the result that the peak fuel processor efficiency would be 23–5% with a hydrogen concentration of 9% and 2%, respectively. Note that the conservative equilibrium temperatures also reflect that true equilibrium compositions are not achieved within the size of the reactor particularly as it is desirable to limit the size of the reactors to reduce the required start-up energy. Achieving ideal performance is also limited due to temperature and mixture non-uniformities within the reactors. Further, due to the low efficiency without water addition, sustained operation without water addition would lead to over temperature without extensive heat removal. Additional heat exchangers would exacerbate the problem of start-up energy. Also, for high overall ATR o/c levels, the air would need to be introduced in stages with intermediate heat removal to avoid excessive reaction temperatures, and the first stage would be prone to soot formation at the lower o/c levels of the first stage. An option for non-water operation is possible by using exhaust gas re-circulation as disclosed by Goebel and Pettit [3]. The dilution by the exhaust gas prevents excessive temperature for the high o/c operation and also provides product water from the fuel cell reaction. Most approaches to fuel processor start-up have been concerned with heating the various components to their operational temperature. However, it is possible to achieve operation before fully heating all of the components. In particular, it is possible to achieve operation by heating a portion of the reactors, but it is also required that steam be generated. Steam is normally generated by heat exchangers in the PrOx reactor or in the combustor exhaust. However, heat exchanger devices are typically more massive than the monolithic supported catalytic reactors for the same surface area due to the double wall construction to isolate the second fluid and the metal material typically used for heat exchangers. Further, these heat exchangers are typically placed downstream of the reactor, so their heating is delayed until the upstream reactor is nearly fully heated. So to achieve rapid start, an alternate source of steam is required during the initial phase of operation until the vaporizers have been heated. Then, for normal operation, it is more efficient to integrate heat from the process for reactant pre-heat or steam generation. The main difference of the fuel processor presented here from other fuel processors is the use of burners with downstream water spray to generate the steam needed for fuel processing to allow rapid fuel processor start-up. Additional options for this approach have been described by Goebel [4]. A risk of this approach could be damage to the catalysts due to water condensation; however, for any fast starting approach, water condensation must be tolerated. While the steam is necessary for fuel processing, it has the additional benefit of increasing the mass flow to heat the reactors. Steam is particularly effective, as it has a higher specific heat than air.
955
3. Experimental A fuel processor was constructed which incorporated two burners with direct steam generation. A schematic diagram is shown in Fig. 2 that has the same orientation as the photograph of the fuel processor in Fig. 3 with the first and second burners in the lower right and upper left corners, respectively. Flanges were used throughout the fuel processor so that any of the catalysts could be replaced. Also, the fuel processor was insulated, but the insulation was removed in the photograph. The layout of the burners is shown in Fig. 4. The fuel processor consisted of the conventional chain of reactors: ATR, WGS and PrOx, that were constructed from catalyst coated monoliths. A brief description of the components proceeding along the flow path follows. The first of the two burners was directed into the fuel processor inlet. The inlet used a solid-cone, direct-injection fuel injector that sprayed fuel co-axially with an air and steam flow. After a thin section of foam to isolate the ATR from the inlet, the first portion of the ATR consisted of conical shaped, catalyzed foam to diffuse the flow from the smaller inlet diameter followed by straight section of catalyzed monolith. A tube-in-shell heat exchanger was used to pre-heat the air and steam going to the inlet in counter-flow to the ATR outlet to cool the reformate before the first WGS reactor (WGS 1). A turn was located between the WGS stages, so a diffuser was required to redistribute the flow into the downstream WGS stage (WGS 2). A water injector was placed in the turn between the stages to provide a temperature reduction. The second burner was also tied into this location as the space provided some mixing between the burner exhaust and the flow from upstream. This second burner was used to directly supply steam to the second WGS and enable the fastest possible start time. Both WGS reactors were based on Platinum group metals rather than CuZn for improved temperature tolerance (especially for the first stage) and oxygen exposure tolerance to allow lean start methods. A two-stage, adiabatic PrOx reactor was used. Each PrOx stage included an array of tubes with holes for distributing the PrOx air followed by a vaporizer to cool the reformate and generate steam from the water supply, and a section of PrOx catalyst. A small heat exchanger in the reformate exit stream from the fuel processor pre-heated the water. This heat exchanger would also cool the reformate which would be needed before a membrane based fuel cell. The steam generated by the PrOx vaporizers was directed through a pressure regulator to the tube-in-shell heat exchanger mentioned previously along with the ATR air. Several drains were placed at key locations throughout the fuel processor to allow condensed water to be periodically drained during the start-up process by actuating the solenoid drain valves. Note that this fuel processor did not use a combustor that would normally burn the anode exhaust from a fuel cell. From a heat integration perspective, the combustor heat input is not required, but if available, could be used to generate additional steam.
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air
water
burn 2
fuel
air
WGS2
water
WGS 1
vap
HX
air
PrOx
ATR
vap
inlet
PrOx
fuel
HX
reformate
water
H2 air
water
burn 1
fuel
air
Fig. 2. Schematic of fuel processor.
Fig. 3. Photograph of fuel processor.
Each burner, as shown in Fig. 4, had an air line, a fuel nozzle, a spark plug for initiating the flame, and a water injector. The air was introduced through a swirler to stabilize the flame. A pressure atomizer was used to spray the fuel, and an injector solenoid was used to regulate the fuel flow. A similar arrangement was used to spray the water into the downstream section of the burner assembly. A conical section separated the flame zone from the water spray. This convergence ensured that the recirculation zone established by the swirler was closed and isolated the water spray from the flame. The cooled exhaust then flowed back over the burner wall to cool it from the flame before exiting. The burners, including the water vaporization sections, were 92 mm in diameter and 235 mm long. They were operated up to 0.43, 10 and 5.2 g/s of fuel, air and water, respectively. This represents a heat input rate of 18.6 kW.
To obtain the fastest possible start, hydrogen was used to ensure catalyst light-off. This light-off hydrogen was introduced into the line feeding the fuel processor inlet. An effort during the course of the testing was to use a minimal amount of hydrogen. Electrically heated catalyst sections could also be used to achieve catalyst light-off. A summary of the full power design conditions is provided in Table 1. The mass of the components is listed in Table 2. The diameter of the reactors was 144 mm except for the ATR where an 80 mm diameter monolith was initially used. In production, the flanges would not be required, so the total mass would be reduced to 26.7 kg. The reactor shells (the metal case and fibrous pad around the reactor monoliths) were also listed separately as this mass may not need to be fully heated to start the fuel processor. The external insulation mass was not included. It should be noted
to fuel processor
S.G. Goebel et al. / International Journal of Hydrogen Energy 30 (2005) 953 – 962
957
er
irl
sw wa sp ter ray
air supply
water injector
flame zone
fuel injector
water drain
spa plu rk g
Fig. 4. Schematic of start-up burner.
Table 1 Full power design conditions
Table 2 Component masses
Flows (g/s) Fuel ATR air PrOx air Water (PrOx) Water (WGS - inj)
2.0 8.2 1.4 4.6 1.0
Component
Ratios o/c (ATR) s/c (ATR) s/c (WGS 2) CO (mole%, WGS1in) CO (mole%, WGS2in) CO (mole%, WGS2out) O2/CO (PrOx)
0.83 1.8 2.2 9.3 3.7 0.6 2.0
Mass (kg)
Reactors Heat exchangers Inlet and mixing sections Reactor shells Manifolds and plumbing Burners Flanges Total
34.1
95 eff eff (CH4 cor) eff (CH4&NMHC cor)
90
68 78
that the total platinum group metal loading was 44.5 g with over three-quarters of this being used in the WGS reactors, so conversion to a CuZn WGS catalyst at least for the larger second stage could significantly reduce the precious metal content.
4. Results and discussion The fuel processor was initially evaluated under steady state conditions, and the results are shown in Fig. 5. The power and efficiency were calculated as outlined below. The power was based on the lower heating value of the hydro-
85 Efficiency (%)
Power, efficiency H2 (kW) Efficiency (%)
5.3 9.6 2.0 3.6 2.6 3.6 7.4
80 75 70 65 60 55 50 0
10
20
30 40 50 Power (kW H2)
60
70
Fig. 5. Efficiency over operating range.
gen from the fuel processor and the efficiency was normalized by the lower heating value of the input fuel. The fuel type used was a mixture of pure components to represent
S.G. Goebel et al. / International Journal of Hydrogen Energy 30 (2005) 953 – 962
1000
1.00 PrOx CO T ATR T WGS1 T WGS2 T PrOx1 T PrOx2 WGS2 CO
900
Temp (C), PrOx (ppm)
800 700 T ATR
600
0.90 0.80 0.70 0.60
WGS2 CO
500
0.50
400
0.40 T WGS1
T WGS2
300
WGS2 CO (%)
958
0.30 T PrOx1
200
0.20 T PrOx2
100
0.10 PrOx CO
0 0
20
40
60
80 100 time (s)
120
140
160
0.00 180
Fig. 6. Start-up temperatures and CO for test to half power.
70
power (kW H2)
60 50
h = (2 − (o/c) − 4a/x − 3a /x − 3b/x)x + y/2 (moles H2 per mole fuel assuming NMHC on C1 basis as Cx H2x with an O2/CO ratio of 1.5 for additional CO clean-up), (2)
Start Energy to Half Power hydrogen input = 7.5 g (1.8 MJ gasoline equiv. at 50% conv. and storage) gasoline input = 136 g (5.9 MJ) hydrogen output = 1.7 MJ (2.3 MJ gasoline equiv. at 75% reforming) start energy = 5.4 MJ (124 g or 0.7 cup) thermal mass (non-flanged) = 5.2 MJ WGS2 air on (to keep warm)
40 run fuel on (over-shoot) hydrogen off (when ATR hot enough for light-off)
30
burner 1 off (when steam from vaporizers)
20 10
burner 2 off (when steam from burner 1 to WGS2)
hydrogen and burner heating
0 0
20
40
60
80 100 time (s)
120
140
160
180
Fig. 7. Start-up power and transitions for test to half power.
a/x = [CH4 ]/([CO2 ] + [CH4] + [NMHC] + [CO]) (from GC measurements),
(3)
a /x = [NHMC]/([CO2 ] + [CH4] + [NMHC] + [CO]) (for NMHC on C1 basis from GC measurements), (4) b/x = [CO]/([CO2] + [CH4] + [NMHC] + [CO]) (from CO gas analyzer and GC measurements), (5)
the range of hydrocarbons found in gasoline. The fuel average composition was Cx Hy where x = 7.494 and y = 14.53 for a molecular weight of 104.5, and had a lower heating value of 4536 kJ/g mole. Metered flows were used for these calculations and corrections were made for any methane, non-methane hydrocarbon (NMHC), or carbon monoxide slip based on gas chromatograph (GC) and CO gas analyzer measurements (the a, a and b terms, respectively). For this fuel processor, as a combustor was not used, the anode stoichiometry did not need to be considered nor a compensation for loss in efficiency as with other fuel processors. Cx Hy + (o/c)x/2 O2 + (o/c)x/23.76 N2 + (s/c)xH2 O → (x − a − a − b)CO2 + h H2 + w H2 O + 0 O2 + a CH4 + a NMHC + b CO + (o/c)x/2 ∗ 3.76 N2 (chemical equation), (1)
efficiency(%) = h LHVH2 (kJ/g mole)/ LHV fuel(kJ/g mole) ∗ 100
(6)
for x = 7.494, y = 14.53, LHV H2 = 242 kJ/g mole, LHV fuel = 4536 kJ/g mole, efficiency(%) = 118.7 − 40 (o/c) (for a, a and b = 0), (7) power (kW H2 ) = efficiency(%)/100 ∗ fuel(g/s) ∗ LHVfuel/MWfuel,
(8)
o/c = air(g/s)/fuel(g/s) ∗ (2 ∗ 0.21/MWair) /(x/MWfuel),
(9)
s/c = water(g/s)/fuel(g/s) ∗ (1/MWwater) /(x/MWfuel)
(10)
S.G. Goebel et al. / International Journal of Hydrogen Energy 30 (2005) 953 – 962
1000
1.00 PrOx CO T ATR T WGS1 T WGS2 T PrOx1 T PrOx2 WGS2 CO
800 700
0.90 0.80 0.70
600
0.60
500
0.50 WGS2 CO
400
0.40
T WGS1 T WGS2
300
WGS2 CO (%)
T ATR
900 Temp (C), PrOx CO (ppm)
959
0.30
T PrOx2
200
0.20 T PrOx1
100
0.10 PrOx CO
0 0
20
40
60
80 100 time (s)
120
140
160
0.00 180
Fig. 8. Start-up temperatures and CO for test to full power.
for MWfuel = 104.5, MWair = 28.85, MWwater = 18, o/c = 0.203 air(g/s)/fuel(g/s),
(11)
s/c = 0.774 water(g/s)/fuel(g/s).
(12)
The efficiency data in Fig. 5 represents the entire range of fuel processor performance. The methane and NMHC slip from this fuel processor were excessively high leading to substantial losses in overall efficiency and power. If this problem were fixed, the efficiency would be close to the expected 80% value. It should be noted that some of the uncorrected efficiencies were high (almost 90%) due to low o/c operation to limit the temperature of hot regions in the ATR (which allowed for hydrocarbon slip in the cool regions). Much lower levels of methane and no NMHC slip were observed on previous tests with similar hardware. Additional evaluations and modifications were undertaken to correct this problem. Temperature profiles across the ATR did indicate non-uniformity under many conditions. At lower power, the inlet exhibited hollow cone injector characteristics with hotter temperatures at the center, and the spray may not have fully penetrated at high power leading to a more uniform temperature (or hotter temperatures at the wall in some cases). Additional 80◦ angle spray injectors were tried, but they continued to follow the previously observed patterns. A 60◦ angle spray injector was also tested and provided more uniform temperature profiles at low power but became un-runnable at higher powers due to high temperatures at the wall. So the 80◦ injector was used for the remainder of the runs. To facilitate mixing, additional catalyzed foam was added to the ATR and a larger monolith section was also used to improve fuel conversion. The replaced ATR catalyst significantly reduced the NMHC slip initially, but after a few tests, the high levels of NMHC slip returned.
Even under conditions where the temperature profile (inlet mixing) was uniform, high levels of hydrocarbon slip were still observed. From the GC speciation, it was noted that toluene (one of the parent molecules) contributed over 50% of the NMHC (on a C1 basis) while octane (the most abundant parent molecule) was not observed. This may indicate a preference of fuel type for fuel processing or poorer catalyst performance on certain types of fuel compounds. Another factor is the fuel pulsation from the injector (which would require storage capacity on the catalyst), however, at full power, the injector was held open. Other notable performance metrics were the PrOx performance and pressure drop. The PrOx reactor performed better than expected reducing CO levels to below 100 ppm with an O2/CO ratio of about 1.5 that was lower than the expected ratio of 2.0 given in Table 1. The pressure drop of the entire fuel processor was only 16 kPa at full power operation at a back-pressure of 50 kPa g. Figs. 6–9 show fuel processor starts using hydrogen for catalyst light-off. Shown are CO levels out of the PrOx and WGS 2 (labeled as PrOx CO and WGS 2 CO, respectively). Also shown are temperatures at the front of reactors. Hydrogen output versus time was calculated from the flows, and this information, as well as fuel input, was used to determine start energy. These charts also indicate the steps in the start sequence. The process was initiated by starting the air flows to Burner 1 and Burner 2 as well as the two PrOx reactors of 10, 5, 1.2 and 0.7 g/s, respectively. So the start-up air flow was somewhat greater than the normal operation air flow. Then the burner fuels were started and the burners were lit. After the burners were lit, the burner water sprays were turned on. The light-off hydrogen was started after burner operation was established. The primary function of the light-off hydrogen was to quickly get the
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S.G. Goebel et al. / International Journal of Hydrogen Energy 30 (2005) 953 – 962 70
50 40 30
run fuel on (over-shoot) hydrogen off (when ATR hot enough for light-off)
20 10 0 0
burner 1 off (when steam from vaporizers)
burner 2 off (when steam from burner 1 to WGS2)
hydrogen and burner heating
power (kW H2)
60
Start Energy to Full Power hydrogen input = 3.5 g (0.9 MJ gasoline equiv. at 50% conv. and full power storage) (when WGS1 gasoline input = 172 g (7.5 MJ) heated) hydrogen output = 3.3 MJ (4.4 MJ gasoline equiv. at 75% reforming) start energy = 3.9 MJ (90 g or 1/2 cup) thermal mass (non-flanged) = 5.2 MJ
20
40
60
80
100
120
140
160
180
time (s)
Fig. 9. Start-up power and transitions for test to full power.
front of the ATR to its light-off temperature, but the lightoff hydrogen was also important to initiate PrOx heating as these reactors were not directly heated by a burner. After the front of the ATR was above its light-off temperature (which was achieved in about 20 s after the start of the process for the test shown in Figs. 8 and 9), the hydrogen was shut off and the run fuel was started and the fuel processor begins to make reformate. It is desirable to use as little start-up hydrogen as possible to reduce the start-up energy requirements. The run fuel was set at a level commensurate with half power under normal operation, but due to the higher o/c operation during start-up, about 30% of full power hydrogen was generated. When the steam from Burner 1 was established through the fuel processor, Burner 2 and its water spray could be shut off. It is desirable to shut off the burners as soon as possible to reduce the start-up energy requirements. A reduced air flow was continued through Burner 2 to allow direct reformate-air reaction on WGS 2 to continue heating this reactor that would otherwise be cooled down as the upstream flow from WGS 1 was cooler at this point. When the PrOx vaporizers began to warm, water flow to these vaporizers was started. When vaporizers began to fully generate steam, Burner 1 and its water spray were shut off. Air for the ATR reaction was then supplied via the normal route through the heat exchanger rather than by the lean exhaust from Burner 1. At this point additional heating of WGS 1 was desired before transitioning to full power. Most of the starts showed a CO spike when the run fuel was started as the fuel flow typically overshot the desired value as the controls adjusted to the fuel pressure. CO spikes were often observed when Burner 1 was shut off as all the steam generation must be supported by the PrOx vaporizers. CO spikes were also observed upon the transition to full power. Fig. 6 shows a start with a longer hydrogen light-off period that had smaller CO spikes than other starts. To achieve low CO and a fast start to full power requires a good sequence of burner operation to sustain the needed steam (and not stopping burner steam generation before the PrOx vaporizers are produc-
ing adequate steam), appropriate air flows to sustain the reactor temperatures (without an over-temperature) and purging to remove condensed water. For the start shown in Fig. 8, the later portion had low CO, so in combination with the first portion of the start shown in Fig. 6, it appears possible to achieve a fast start with low CO throughout. From fuel cell testing, brief spikes over 1000 ppm of CO are tolerable as long as the average CO level can be handled. A drop in ATR temperature shows the need to purge excess water during the start sequence (this occurs when burner 1 is shut off as air then comes through the ATR HX which probably has condensed water in the bottom of the shell). These start-up sequences were performed using manually initiated steps. Additional controls development could improve the consistency of the start-up process. To put the start energy into perspective, it is compared to the energy used for driving. For example, consider a reformate based fuel cell vehicle that achieves 60 mpg during driving. It would require 30.7 MJ of fuel over a 15 mile drive cycle. For the fuel processor start energy, the fuel input, start-up hydrogen and additional air compressor work are considered. Fuel input for start-up was taken to be the amount above what was reformed. The amount of fuel reformed was based on the amount of hydrogen produced using a 75% conversion efficiency. The start-up hydrogen was included at twice the energy (to account for formation and storage efficiency) and some compression work (about 0.3 MJ on a fuel equivalent basis) was included to account for the additional air flow during start-up. On this basis, the fuel processor would require a fuel equivalent start energy of 4.2 MJ (for the start-up shown in Fig. 9). Compared to the fuel used for driving in this example, this is an additional 14% fuel. Including this additional start-up energy in the total drive cycle fuel consumption, the drive cycle average fuel economy would be reduced to 53 mpg. The thermal portion of the start energy of 3.5 MJ is significantly less than the thermal energy of 5.2 MJ to heat the entire fuel processor (excluding the flanges and burners) to operating conditions which indicates that the fuel processor can operate at full power conditions without being fully heated (probably the reactor shells are not fully heated). This also indicates that some of the normal operation waste heat (which typically appears as steam in the reformate output) can be used to heat the fuel processor during start-up. This also demonstrates that the energy used for this start-up process was effectively utilized. To store the start-up hydrogen, using a 5000 psig (341 MPa) tank, only 0.012 l per start would be required. Such a tank could be replaced at each fuel refill. Alternative start methods without using hydrogen were also investigated. One approach involves generating reformate by operating a burner fuel rich and requires that the catalysts light-off on this reformate. Reformate light-off capability was evaluated by establishing air flows and then Burner 1 was lit under fuel rich conditions (with water spray to control the burner temperature). Burner 1 was used to
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1000
1.00 PrOx CO T ATR T WGS1 T WGS2 T PrOx1 T PrOx2 WGS2 CO
800 700 600
0.90 0.80 0.70 0.60
T ATR
500
0.50
WGS2 CO
400
0.40
T WGS2
300
WGS2 CO (%)
900 Temp (C), PrOx CO (ppm)
961
0.30 T PrOx1
200
0.20
T PrOx2
100
T WGS1
0.10
PrOx CO
0
0.00 0
20
40
60
80
100 120 140 160 180 200 220 240 time (s)
Fig. 10. Start-up temperatures and CO for lean start.
70
power (kW H2)
60 50 40
Start Energy to Full Power hydrogen input = 0 g (non hydrogen start) gasoline input = 202 g (8.8 MJ) hydrogen output = 3.3 MJ (4.4 MJ gasoline equiv. at 75% reforming) start energy = 4.4 MJ (101 g or 0.6 cup) brief fuel increase thermal mass (non-flanged) = 5.6 MJ (to moderate ATR temperature) run fuel on (over-shoot) (when ATR hot enough for light-off)
30 20
full power (when WGS1 heated)
burner 2 and 1 off (when steam from vaporizers)
10 burner heating
0 0
20
40
60
80 100 120 140 160 180 200 220 240 time (s)
Fig. 11. Start-up power and transitions for lean start.
avoid any direct heating of the catalysts and to delay the arrival of the condensation wave that appeared as a temperature rise in WGS 2 about 15–20 s after the burner was lit. It was found that the WGS catalyst did and the PrOx catalyst did not light-off on fuel rich burner exhaust and air. Only after an inlet temperature rise did significant reactions begin to occur on the PrOx catalyst. Another concern with this approach was generation of soot by the fuel rich flame. Examination of the hardware after a few tests did show soot deposits on the reactors. Based on the PrOx catalyst not lighting-off and burner sooting, it was decided not to pursue fuel rich burner starts. Non-hydrogen start-ups were successfully done by operating both burners fuel lean as shown in Figs. 10 and 11. Here, the burners function to warm the reactors to their lightoff temperatures and also provide steam until the vaporizers are heated. The water spray also reduces the burner exhaust
temperature and increases the heating mass flow. The process was initiated by starting the air flows to Burner 1 and Burner 2 of 10 and 5 g/s, respectively. Then the burner fuels were started and the burners were lit. After the burners were lit, the burner water sprays were turned on. After the front of the ATR was above its light-off temperature (which was achieved about 50 s after the start of the process for the test shown in Figs. 10 and 11), the run fuel and PrOx air flows were started and the fuel processor begins to make reformate. The run fuel was set at a level somewhat below half power under normal operation to operate at a slightly higher o/c to assist heating of the ATR. When the PrOx vaporizers began to warm, water flow to these vaporizers was started. When the vaporizers began to generate steam, Burners 1 and 2 were shut off. The CO spike observed at this point may be due to the steam generation of the PrOx vaporizers being at a rate lower than that produced by the burners with water spray. Air for the ATR reaction was then supplied via the normal route through the heat exchanger rather than by the lean exhaust from Burner 1. A reduced air flow was continued through Burner 2 to allow direct reformate-air reaction on WGS 2 to continue heating this reactor that would otherwise be cooled down as the upstream flow from WGS 1 was cooler at this point. Additional heating of WGS 1 was desired before transitioning to full power. WGS 1 was slow to heat, so it may have been beneficial to extend the duration of Burner 1 operation.
5. Conclusions The criticality of steam to the fuel reforming process was illustrated. By utilizing direct vaporization of water in burner
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exhaust, and hydrogen for catalyst light-off, excellent start performance was obtained with a start time of 20 s to 30% power and 140 s to full power. The requirement to have water for a fuel processor start (and operation) clearly creates problems for automotive applications that must tolerate subfreezing conditions. Start methods not using light-off hydrogen were also tested. Issues with PrOx light-off and burner sooting were found with the fuel rich start method. A fuel lean start method to full power was demonstrated with 50 s to start and 190 s to full power. Again the use of water spray was beneficial to increase the heating mass flow and necessary to provide steam until the vaporizers were heated.
Acknowledgements A much larger team was working on fuel processing at General Motors, and they provided the catalysts, heat exchangers and build support for this fuel processor. The support of General Motors and its management team for the opportunity to demonstrate the capabilities of fuel processing for automotive fuel cell systems was also greatly appreciated.
References [1] Robb GM, Pettit WH. Method for quick start-up of a fuel processing system using controlled staged oxidation. US Patent Appl. US2003/0170510; 2003. [2] Goebel SG, Pettit WH, Sennoun MEH, Miller DP. Staged lean combustion for rapid start of a fuel processor. US Patent Appl. US2003/0093949; 2003. [3] Goebel SG, Pettit WH. Fuel processing system having gas recirculation for transient operations. US Patent Appl. US2003/0136880; 2003. [4] Goebel SG, Direct water vaporization for fuel processor startup and transients. US Patent Appl. 2003/0154654; 2003.