Catalytic Exhaust Gas Recirculation-Loop Reforming for High

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Catalytic EGR-Loop Reforming for High Efficiency in a Stoichiometric SI Engine through TCR and Dilution Limit Extension. 1. Catalyst Performance and Fuel Effects Yan Chang, James P. Szybist, Josh A Pihl, and Daniel William Brookshear Energy Fuels, Just Accepted Manuscript • DOI: 10.1021/acs.energyfuels.7b02564 • Publication Date (Web): 19 Dec 2017 Downloaded from http://pubs.acs.org on December 30, 2017

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Catalytic EGR-Loop Reforming for High Efficiency in a Stoichiometric SI Engine through TCR and Dilution Limit Extension. 1. Catalyst Performance and Fuel Effects Yan Chang, James. P. Szybist,* Josh A. Pihl, and D. William Brookshear Fuels, Engines, and Emissions Research Center, Oak Ridge National Laboratory, NTRC Building, 2360 Cherahala Blvd, Knoxville, Tennessee, 37932, United States KEYWORDS: Reforming, EGR, hydrogen, thermochemical recuperation ABSTRACT: The use of fuel reformate from catalytic processes is known to have beneficial effects on the spark-ignited (SI) combustion process through enhanced dilution tolerance and decreased combustion duration, but in many cases reformate generation can incur a significant fuel penalty. In this two-part investigation, we demonstrate that efficient catalytic fuel reforming can result in improved brake engine efficiency while maintaining stoichiometric exhaust under the right conditions. In part one of this investigation, we used a combination of thermodynamic equilibrium calculations and experimental fuel catalytic reforming measurements on an engine to characterize the best possible reforming performance and energetics over a range of equivalence ratios and O2 concentrations.

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Ideally, one might expect the highest levels of thermochemical recuperation for the highest catalyst equivalence ratios. However, reforming under these conditions is highly endothermic, and the available enthalpy for reforming is constrained. Thus for relatively high equivalence ratios, more methane and less H2 and CO are produced. Our experiments revealed that this suppression of H2 and CO could be countered by adding small amounts of O2, yielding as much as 15 vol % H2 at the catalyst outlet for 4 < Φcatalyst < 7 under quasi-steady-state conditions. Under these conditions the H2 and CO yields were highest and there was significant water consumption, confirming the presence of steam reforming reactions. Analyses of the experimental catalyst measurements indicated the possibility of both endothermic and exothermic reaction stages and global reaction rates sufficient to enable the utilization of higher space velocities than those employed in our experiments. In a companion paper detailing part two of this investigation, we present results for the engine dilution tolerance and brake engine efficiency impacts of the reforming levels achieved. INTRODUCTION Efficiency improvements in spark-ignited engines are desired as a way to meet increasingly challenging fuel economy and CO2 emission regulations.1 External cooled exhaust gas recirculation (EGR) provides known thermodynamic benefits while maintaining compatibility with conventional three-way catalysts for emissions control.2 Caton3 summarized the thermodynamic benefits of EGR, which include reduced pumping work at part-load conditions, decreased heat transfer due to lower cylinder temperature, and increased ratio of specific heats. External EGR is also a proven way to decrease the knocking propensity for a given fuel, which can be used as the basis for additional increases in efficiency through more advanced combustion


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phasing or higher compression ratio.4 Lastly, EGR reduces NOx emissions over a broad range of speed and load conditions.5 The amount of EGR dilution that can be used is limited due to cycle-to-cycle combustion instability, thereby limiting the potential efficiency benefit of EGR.6-8 The root cause of the cyclic instability with EGR has been linked to a decrease in flame speed, which elongates the initial flame kernel development process, making it more susceptible to stochastic turbulence variation.9-11 While several technologies unrelated to fuel reforming are being developed to extend the EGR dilution limit, including high-energy long-duration ignition systems12 and incorporation of different higher turbulence combustion chamber flows,13 the focus of this work is to extend the EGR dilution limit by using the reformed products of the fuel to increase the EGR dilution tolerance. The addition of the major products of fuel reforming, H2 in particular, increases the combustion rate and can increase engine efficiency because of the shorter combustion duration.14-20 Alger et al.21 reported that H2 can also be used to extend the EGR dilution limit, with the addition of 1 vol % H2 extending the EGR limit from 25% to more than 50% for gasoline and from 20% to 28% for compressed natural gas. Fennell et al.22 showed that simulated H2-rich reformate could extend EGR dilution in an SI engine from 21% to 27% at the same combustion stability. Ivanič et al.23 added reformate at levels of 15% and 30% gasoline energy equivalent in a single cylinder engine and found that, under part-load conditions, lean dilution can improve engine efficiency by as much as 12% while EGR dilution delivers 8% improvement. Though there are many approaches to generating reformate onboard, they can generally be classified into two broad categories. The first category is where fuel is reformed in one or more


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cylinders in an engine using noncatalytic processes. This category includes the Dedicated EGR (D-EGR) strategy developed by Southwest Research Institute, which is the most developed strategy in this category.24 D-EGR uses fuel-rich combustion in one cylinder and recirculates its exhaust to the intake system, generating brake thermal efficiency as high as 42.5%25 and demonstrating a vehicle-level fuel consumption decrease of more than 10%.26 This category also includes injecting fuel during the negative valve overlap period for a homogeneous charge compression ignition engine to manipulate the fuel-air mixture autoignition propensity.27-32 The second category of fuel reforming, and the category of the present work, is where a catalyst is used to reform the fuel outside of the engine cylinders19,22,33-40 Catalytic reforming processes often use the sensible enthalpy in the exhaust to promote reforming reactions over the catalyst to improve the energetics of reforming.18,19,33-36,41-43 The range of possible combustion and reforming reactions that can occur has been presented by Ahmed and Krumpelt44 and by Jamal and Wyszynski,18 and is presented below for completeness using iso-octane as the starting fuel. Equation 1 shows the reaction stoichiometry for the complete combustion of iso-octane, while Equation 2 defines equivalence ratio (Φ). For the purposes of this manuscript, Φ refers to the O2/Fuel ratio of complete combustion in Equation 1 relative to the actual O2/Fuel ratio. Equations 3–5 show the primary types of reactions that can occur in a reformer. When O2 is present such that the molar O/C ratio is unity, partial oxidation (POx) occurs where the parent hydrocarbon is converted to CO and H2, resulting in an exothermic process as shown in eq 3. For an environment devoid of O2 but where H2O is present, the resultant steam reforming reaction, shown in eq 4, is highly endothermic and results in much higher concentrations of H2. Similarly, eq 5 shows a dry reforming reaction where CO2 is consumed to form H2 and CO in a process that is substantially more endothermic than steam


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reforming, but where the concentrations of H2 are lower. Ideally, the endothermic steam reforming reactions are driven by the sensible enthalpy in the exhaust, allowing for waste heat recovery through thermochemical recuperation (TCR).45 In this work, we studied reforming environments that contain an insufficient amount of O2 to convert all of the fuel to CO and H2 through eq 3 and thus consist of a balance of POx, steam, and dry reforming processes. Complete Combustion: C8H18 + 12.5 O2 → 8 CO2 + 9 H2O (- 5050.2 kJ/mol)

Equivalence Ratio: Φ

=

(1)

(2)

POx Reforming: C8H18 + 4 O2 → 8 CO + 9 H2 (−660.4 kJ/mol)

(3)

Steam Reforming: C8H18 + 8 H2O → 8 CO + 17 H2 (1,274.2 kJ/mol)

(4)

Dry Reforming: C8H18 + 8 CO2 → 16 CO + 9 H2 (1,603.8 kJ/mol)

(5)

The work by Leung et al.38 showed that, for ethanol, the steam reforming reaction (eq 4) is very active at engine exhaust temperatures, converting nearly 100% of the fuel energy at 600°C. In contrast, dry reforming (eq 5) produces less than 40% fuel conversion at the same temperature. This indicates that dry reforming is not the favored reaction pathway, likely due to the larger energy input required, especially at lower initial temperatures. Regardless of whether reforming takes place via the steam or dry reforming pathway, though, the product concentrations can equilibrate via the water-gas shift (WGS) reaction, shown in eq 6. WGS: CO+H2O→CO2+H2 (−41 kJ/mol)

(6)


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It can be challenging to operate the engine and catalyst systems together so that good performance of the catalytic reformer and engine are achieved simultaneously. Hwang et al.36 integrated a reforming catalyst containing Rh and Pt into the exhaust manifold of a diesel engine. While they were able to produce high concentrations of H2 (> 10%), the reforming system resulted in an overall engine efficiency decrease because the reformer catalyst equivalence ratio was too low (Φcatalyst as low as 1.5), and the reforming process was too dependent on the exothermic reactions in eq 3. This is consistent with much of the diesel reforming work where high concentrations of H2 can be produced, but because of the dependency on POx reforming, there is a large fuel energy penalty to produce the H2.46,47 In contrast, Ashida et al.33 used a steam reforming catalyst (4 wt% Rh/Al2O3, La additive) in the EGR loop of a spark ignition engine without any additional O2 (Φcatalyst approaching ∞). They initially found high levels of fuel conversion and H2 production, but the steam reforming catalyst began deactivating immediately, with a 35 ppm S fuel resulting in a 90% deactivation within 5 hours. In this manuscript, we characterize reformer catalyst performance over a range of Φcatalyst conditions in an effort to find a balance between the POx reactions required for sustained catalyst activity and durability and the endothermic steam reforming reactions required to minimize energy losses or to achieve TCR during reforming. This experimental study was conducted using a full-sized catalyst on a multi-cylinder engine but builds on previous bench flow reactor experiments conducted by Brookshear et al.48 This paper characterizes reformer performance and energetics using iso-octane with the EGR-loop reforming catalyst under a variety of operating conditions while a companion paper49 focuses on whole-engine performance, including EGR dilution tolerance, emissions, and brake thermal efficiency. METHODOLOGY


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Equilibrium Calculations We determined the limiting chemical equilibrium products for a range of conditions in the reforming reactor using the equilibrium solver in ANSYS Chemkin-Pro 17.2. All of the 1,389 species in the Lawrence Livermore National Laboratory (LLNL) gasoline surrogate kinetic mechanism (and their associated thermodynamic properties) were included, allowing the resulting products to be fully unconstrained.50 Solid carbon or graphite was not included as a possible product in this reaction mechanism. A range of equivalence ratios was included, with the as-fed CO2, O2, H2O, and N2, and the parent fuel at varying concentrations. Iso-octane (H/C = 2.25) was used for the fuel, and the calculations were performed for initial O2 mole fractions ranging from 0.005 to 0.06 and for equivalence ratios from 3.0 to 10.0. Equilibrium calculations were conducted at a constant pressure of 1 atm and initial temperatures of 500°C and 700°C. Experimental Facility The engine used in this study was a 2.0 L GM Ecotec LNF SI gasoline engine equipped with the production side-mounted direct injection fueling system. Engine geometry details are presented in Table 1. The combustion chamber geometry and camshaft profiles were unchanged from the stock configuration. All engine experiments presented in this manuscript were conducted with primary reference fuel iso-octane obtained from Haltermann, as shown in Table 2. This manuscript focuses on the catalyst operation, and data presented here represent just one of the four cylinders operating and feeding the catalyst, referred to as cylinder 4. A companion manuscript presents the results for the full engine operation.49 The engine configuration is shown in Figure 1.


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Table 1. Engine geometry. Bore × Stroke [mm] Conrod length [mm] Wrist pin offset toward expansion stroke[mm] Compression ratio [-] Fuel injection system

86.0 × 86.0 145.5 0.8 9.2/1 Direct injection, sidemounted, production injector with opposite linear wall directed sixhole spray pattern

Table 2. Fuel Properties RON (ASTM D2699) MON (ASTM D2700) S (RON–MON) Boiling point C wt % H wt % O wt % LHVa [MJ/kg]

Iso-octane 100 100 0 99.3 84.2 15.8 0 44.3

Figure 1. Schematic of the engine and test cell configuration for the catalyst-only experiments.


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The engine was operated using laboratory fueling and air handling systems. A lift pump delivered fuel at 5 bar pressure through a Coriolis effect flow meter to the engine, where the cam-driven fuel pump was used to maintain a constant rail pressure of 100 bar throughout the experiment. The air for cylinder 4 was supplied using a mass flow controller. A Drivven engine controller with the Combustion Analysis Toolkit package was used to control the engine and acquire crank angle (CA)–resolved data. For each experimental condition, cylinder pressure, spark discharge, and camshaft position data were recorded at 0.2°CA resolution for 1,000 sequentially fired cycles. Cylinder pressure was measured using a flush-mounted piezoelectric pressure transducer from Kistler (6125C), and camshaft position was recorded from the production Hall-effect sensors. Start of fuel injection timing was commanded during the intake stroke and was held constant at 280°CA before firing top dead center (TDC), and the start of post-injection was held constant at 150°CA after TDC. Engine coolant temperature was maintained at 90°C. Concentrations of water, CO2, CO, CH4, and iso-octane were performed using a MultiGas 2030HS Fourier transform infrared (FTIR) instrument from MKS Instruments. The FTIR was calibrated to detect up to 20% water using an 8-point calibration and 23% CO2 using a 14-point calibration. Two different calibrations for CO were, with the low detection range being a 5000 ppm, 15-point calibration, and the high detection range being an 8%, 10point calibration. The calibrated ranges for CH4 was 3000ppm using a 14-point calibration, and for iso-octane, a 1000ppm range using a 5-point calibration was used. Hydrogen was measured using an ODYSSEY magnetic sector mass spectrometer from VTI Instruments. A gas divider was used to verify that there was a linear response up to 10% H2, thereafter a single-point calibration of 10% was used for daily calibration. Smoke emissions were monitored with a reflective filter smoke number instrument, but all smoke measurements registered a negligible


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filter smoke number (< 0.01), thus further discussion of smoke emissions is omitted from this manuscript. Primary reference grade iso-octane, obtained from Haltermann Solutions, was used as the fuel in this investigation. The reforming catalyst was a precommercial 2 wt % Rh supported on Al2O3, which was synthesized and coated on a cordierite monolith substrate by Umicore Autocat USA. Two cordierite catalyst monoliths were used in the experiment, each measuring 106 mm in diameter and 106 mm in depth, providing a total catalyst volume of 1.9 L. The catalyst cell density was 400 cells per square inch. Temperature was monitored at the catalyst inlet, between the two catalyst monoliths, and at the catalyst outlet, also shown in Figure 1.

Experimental Methodology The engine was operated in such a way as to provide the catalyst with a fuel-rich mixture that also contained oxygen. This was accomplished by operating the main combustion event in cylinder 4 fuel-lean and providing a post-injection of fuel near the exhaust valve opening timing. A schematic of the cylinder pressure and the fuel injection schedule for this strategy is shown in Figure 2. The main injection event was held constant throughout this study, delivering 16 mg/injection event. Both the level of excess air and the post-injection fuel mass were variable.


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Figure 2. Cylinder pressure and fuel injection strategy for operation with catalyst reforming. A fuel-lean Φcombustion combined with a post-injection provides a fuel-rich Φcatalyst with oxygen present. We adjusted the main combustion event among four Φcombustion conditions (Φcombustion = 0.91, 0.83, 0.77, and 0.71), resulting in four nominal oxygen flow rates to the catalyst ( = 6.1, 11.9, 17.5, and 23.3 g/min). A portion of the fuel that was delivered in the post-injection was not exhausted and sent to the catalyst, but rather was included in the trapped residuals and consumed in the subsequent combustion cycle. For the purpose of presenting results, that portion of fuel was neglected when calculating Φcombustion, and Φcatalyst. As a practical matter, though, the trapped residual fuel prevented us from precisely knowing the fuel energy and the composition being delivered to the reforming catalyst, effectively preventing elemental and energy balances to be conducted on the experimental results. Modifications are being pursued to the experimental methodology to allow these analyses for follow-on work.

At each oxygen flow rate, the post-injection fuel was varied to yield an estimated Φcatalyst ranging from 3.0 through 10.0. The minimum fuel delivered in the post-injection event was 4.8 g/min, or


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about 10% of the fuel required by the remaining three cylinders to achieve 4.0 bar brake mean effective pressure (BMEP) at 2,000 rpm. The maximum fuel delivered in the post-injection was 66.6 g/min, or about 135% of the fuel required by the remaining three cylinders. Thus, this range of fuel flux over the catalyst, presented graphically in Figure 3, represents a wide operating range (note that the markers for conditions I, II, and III in Figure 3 are referred to in the Results section).

Figure 3. Fuel flow rate through the catalyst as a function of catalyst O2 flow rate and Φcatalyst.

We followed a systematic procedure at each experimental condition to ensure that if coking did occur, the catalyst was regenerated prior to the next condition. This procedure consisted of the following steps. 1. Operating the engine without the post-injection at the desired Φcombustion and allowing the catalyst to thermally stabilize. 2. Once the catalyst stabilized, the data acquisition was started and allowed to run for about 1 min before the post-injection was started.


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3. The post-injection was introduced, causing a step-change in Φcatalyst, and the system was allowed to operate under these conditions for about 7 min. 4. The excess air and the post-injection were stopped nearly simultaneously to produce Φcombustion = 1.0. 5. The Φcombustion was gradually decreased by increasing the excess air to cylinder 4 over the course of 5 min to allow any coke formed on the catalyst to be gradually oxidized, preventing a large and potentially damaging spike in temperature. Throughout the manuscript, the results from these conditions are presented in two ways: (a) the transient response to the step change in Φcatalyst with the post-injection introduction and (b) the quasi-steady-state temperature profiles and species concentrations at the catalyst exit produced by the reforming process. RESULTS AND DISCUSSION ENERGETICS AND EQUILIBRIUM ANALYSIS The energetics of the reforming process are dependent on the relative amount of POx, steam, and dry reforming, the H/C ratio of the fuel, and competing equilibrium products. Equations 3 and 4 introduced the energetic differences of reforming in the introduction for a single fuel. Figure 4 introduces the effect of H/C ratio by showing the ratio of product enthalpy of combustion to reactant enthalpy of combustion for a series of alkanes, olefins, and aromatics. The POx reforming reactions were considered at an atomic C/O ratio of unity, as shown in Equation 3. The individual species enthalpy values were taken from Szybist et al.51 and are referenced at a constant temperature and pressure condition and therefore do not consider sensible enthalpy effects. An enthalpy ratio less than unity represents a chemical energy loss during the reforming


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process, whereas an enthalpy ratio of greater than unity represents a chemical energy increase, or TCR. Figure 4 illustrates the limiting energetics when the reactions are constrained to a single reforming pathway. POx reforming results in a chemical energy loss for all fuels, whereas steam and dry reforming result in a chemical energy increases. The chemical energy loss for POx reforming decreases as the H/C ratio of the fuel increases. For example, methane (H/C = 4) experiences only a 4% energy loss whereas benzene (H/C = 1) experiences a 23% energy loss. The enthalpy ratio of all fuels considered shows a 15% to 25% chemical energy increase for steam reforming, and a 22% to 32% increase for dry reforming. Clearly the H/C ratio is not the only factor affecting the enthalpy ratio. Most notably, all mono-olefins have the same H/C ratio (H/C = 2), but they do not all have the same enthalpy ratio during reforming. It is interesting to note that the olefin with the lowest enthalpy ratio is ethylene, and then the enthalpy ratio monotonically increases as the molecular weight of the mono-olefin increases. This suggests that while the bond energy in the olefin unsaturation is higher, the energy does not get recovered in the reforming process.

Despite the presence of other factors, understanding the reforming

thermodynamics in the context of the H/C ratio is insightful.


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Figure 4. Enthalpy ratio of reforming as a function of the fuel H/C ratio for POx and steam reforming. To understand the more complex realistic equilibrium product distributions, calculations were performed for iso-octane using CHEMKIN over the range of initial temperatures, equivalence ratios, and oxygen concentrations investigated experimentally. It is important to note that equilibrium calculations are the result of the minimization of Gibbs free energy and do not consider the reaction pathways required to achieve the equilibrium condition. Figure 5 shows the equilibrium reforming enthalpy ratios for two different initial temperatures, 500°C and 700°C, presented as functions of Φcatalyst of the mixture and the O2 mole fraction. In addition to Φcatalyst, the x-axis is also labeled with the molar O/C ratio, as this is commonly used in describing reforming stoichiometry. First, it is notable that the enthalpy fraction is significantly higher for the 700°C initial temperature, particularly for Φ > 5. This indicates that the reforming energetics are enthalpy-limited at high Φcatalyst. Steam reforming and dry reforming are highly endothermic and thus require considerable initial sensible enthalpy to drive the equilibrium products to H2 and CO. The enthalpy fraction results fall into two distinct operating


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regimes. For Φcatalyst < 5, the enthalpy ratio is primarily a function of Φ, with very little dependence on the O2 mole fraction. In this region, POx dominates, and fuel energy is consumed during the reforming process. For Φcatalyst > 5, the enthalpy ratio becomes more dependent on the O2 mole fraction, with higher O2 mole fractions resulting in reduced enthalpy ratios. At these conditions, a higher O2 mole fraction corresponds to a higher mole fraction of the fuel to maintain a constant Φcatalyst. As a result, more energy input is needed to drive the endothermic reforming reactions, eq 4 and 5. At lower O2 flow rates, the diluents carry sufficient sensible enthalpy to convert the fuel to H2 and CO. At higher O2 flow rates, the diluents make up a smaller fraction of the total and do not have sufficient sensible enthalpy. Thus, the higher the O2 mole fraction, the higher the demand for energy input to drive the reaction toward the products with high chemical enthalpy: H2 and CO.


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0.06

O2 Mole Fraction

500C

0.05 0.04

1.1

0.03 0.02 0.01 1.0

0.06

O2 Mole Fraction

0.05

700C

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Energy & Fuels

0.9

0.04 0.03 0.02 0.01

0.8

3

4

5

6

7

8

9

10

Catalyst Equivalence Ratio 1

0.8

0.6

0.5

0.4

0.35

O/C Ratio

Iso-octane

Figure 5. Enthalpy fraction of reforming products to the initial fuel with 500°C and 700°C for iso-octane. Regardless of whether the initial reforming reactions are steam reforming (eq 4) or dry reforming (eq 5), WGS reactions will drive the system toward an equilibrium H2/CO ratio. The H2/CO ratio is a strong function of temperature, as shown in Figure 6 for a single starting condition (Φcatalyst = 6, O2 mole fraction = 0.03). Thus, regardless of whether the initial mechanistic step is steam or dry reforming, WGS reactions will favor H2 at low temperatures and favor CO at higher temperatures. It is interesting to note that the equilibrium H2/CO ratio at 400°C is close to the H2/CO ratio from steam reforming alone (H2/CO = 2.125, as shown in eq 4), and the H2/CO ratio


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Energy & Fuels

from dry reforming is asymptotically approached as the temperature exceeds 1400°C (H2/CO = 0.5, eq 5). The catalyst-out temperatures for the experimental portion of this study are all below 625°C. 2.5 Steam H2/CO Ratio

2.0

H2 / CO Ratio

1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 21 22 23 24 25 26 27 28 29 30 31 32 33 34 35 36 37 38 39 40 41 42 43 44 45 46 47 48 49 50 51 52 53 54 55 56 57 58 59 60

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1.5 POx H2/CO Ratio

1.0

0.5

Dry H2/CO Ratio

0.0 400

600

800 1000 Temperature [ C ]

1200

1400

Figure 6. Equilibrium H2/CO ratio as a function of temperature for initial conditions of Φcatalyst = 6, O2 mole fraction = 0.03. There are three major reforming products in the equilibrium calculations: H2, CO, and CH4. The enthalpy fraction of each of these components relative to the reactant enthalpy is shown in Figure 7 for an initial temperature of 700°C. The H2 enthalpy fraction resembles the overall enthalpy fraction shown in Figure 5, exhibiting a maximum coincident with the overall enthalpy fraction maximum. The H2 enthalpy accounts for anywhere from < 25% to > 75% of the initial fuel enthalpy. Relative to the other species, the CO enthalpy fraction shows only a small amount of variability over this range of conditions, accounting for 30%–50% of the total fuel energy. The CH4 formation is highly dependent on the Φcatalyst and oxygen mole fraction conditions. There is very little CH4 formation for a Φcatalyst < 5, but substantial CH4 at a Φcatalyst > 5 at high O2 concentrations. The CH4 formation occurs in the region where the enthalpy balance becomes less


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favorable (Figure 5) and the H2 concentration decreases (Figure 7). Thus, the enthalpy balance is highest when H2 and CO are formed, but CH4 can also be a competing equilibrium product that degrades the enthalpy balance at high O2 mole fraction concentrations.

Figure 7. Enthalpy fraction of H2 to the initial fuel with 700°C initial temperature for iso-octane. Figure 5 shows that the reforming enthalpy ratio is highly dependent on the initial temperature for a Φcatalyst > 5, with higher temperatures resulting in a more favorable enthalpy balance.


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Figure 8 shows that the less favorable enthalpy ratio can be attributed to the formation of CH4. Increasing the initial temperature from 500°C to 700°C in Figure 8 results in a substantial decrease in CH4 in this region, though significant concentrations of CH4 persist.

Figure 8. CH4 equilibrium mole fraction with 500°C (top) and 700°C (bottom) initial temperature for iso-octane. In summary, the equilibrium calculations reveal reforming products should be highly dependent on the initial temperature, Φcatalyst, O2 mole concentration, and H/C ratio of the fuel. The most favorable energetics occur at a high Φcatalyst with low O2 mole fraction so that the diluent gases carry a sufficient sensible enthalpy to drive the endothermic steam reforming processes, thus avoiding the less favorable CH4 formation. However, equilibrium calculations are based solely on thermodynamics and do not account for the actual reaction rates or the selectivity that may


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occur over the catalyst. In an engine environment, the timescale of reaction is limited by the size of catalyst and gas flow rate so that the products can differ from equilibrium results, and the reaction pathways over the catalyst may not reflect equilibrium. Furthermore, there can be substantial gradients in the catalyst temperature under realistic operating conditions, so the favored equilibrium products can vary significantly as a function of position within the catalyst monolith. Nonetheless, an equilibrium calculation allows us to set expectations for the energetics of the reforming reactions. Achieving Quasi-Steady-State Reforming The experimental procedure used to provide fuel to the catalyst was described in the Experimental Methodology section, where the post-injection to provide the fuel to the catalyst is introduced as a step-change. The catalyst does not immediately achieve a steady-state condition, but instead comes to a quasi-steady-state condition over several minutes as the catalyst achieves thermal equilibrium. The path to achieving a quasi-steady-state operating conditions can vary significantly depending on the O2 flow rate and Φcatalyst. Thus, the purpose of this section is to describe the transient response of the catalyst to the fueling step-change in an effort to understand the reforming process in more detail. It should be noted that the transients observed here are not intended to simulate vehicle operation, but rather to provide a better phenomenological understanding. More realistic transient operation will be a focus of future work. The actual reforming catalyst performance was measured at each of the O2 flow rate and Φcatalyst combinations shown in Figure 3. Each operating condition started with lean operation (Φcombustion < 1) to set the O2 flow rate, followed by introduction of the post-injection, causing a step-change


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in the Φcatalyst. With the step-change in Φcatalyst, it takes several minutes for the catalyst to achieve a quasi-steady-state condition. This section highlights the transient response of the catalyst to a step-change in Φcatalyst at conditions I, II, and III, which are indicated in Figure 3. Condition I, shown in Figure 9, is at an O2 catalyst flow rate of 12 g/min and Φcatalyst = 3. When the fuel is introduced via post-injection, there is a slow increase in the catalyst mid-bed and outlet temperature, indicating that this is an exothermic process. The catalyst outlet temperature reaches a steady value after about 4 min of post-injection fueling. The CO2 and H2O concentrations do not change significantly with the introduction of the post-injection. There is a step-change increase in the H2 and CO concentrations, from near zero to about 5% CO and 7%– 8% H2. Additionally, CH4 is formed at a concentration of 1,000 ppm. Nearly all of the input isooctane is converted across the catalyst, with < 200 ppm iso-octane measured at the catalyst output. Note that the iso-octane measurement in this plot is quantified from FTIR spectra, and the ability to differentiate it from other alkanes or alkyl groups heavier than propane is limited. Nonetheless, this trend provides an estimate of unconverted fuel across the reforming catalyst.


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Figure 9. Transient response of the exit concentrations of catalytic reforming products and the monolith temperature profile at an O2 catalyst flow rate of 12 g/min and Φcatalyst = 3 (condition I in Figure 3). Condition II in Figure 3 also uses an O2 catalyst flow rate of 12 g/min but at a richer condition, Φcatalyst = 7, and is shown in Figure 10. In contrast to condition I, where the mid-bed and outlet temperatures increased as a function of time, each of these temperatures decrease by more than 100°C within 2 min of the initiation of the post-injection. A decrease in temperature is indicative of the endothermic steam reforming process. Also indicative of steam reforming is the lower H2O concentration and the higher H2 concentration compared to condition I. The H2O concentration decreases from about 12% before the post-injection to about 5% during the post-


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injection, indicating that it is being consumed in the catalyst as a reactant during steam reforming, and the concentration of H2 is 13%–15%, about double the concentration of condition I. Taken together, the trends associated with CO, CH4, and iso-octane reveal an interesting behavior. There is a spike in CO and CH4 immediately after the post-injection is introduced, with more than 10% CO and 1.5% CH4. The initial spike slowly decreases as the iso-octane concentration at the catalyst outlet increases, while simultaneously the catalyst-out temperature decreases. This demonstrates that as the catalyst temperature decreases, the fuel conversion decreases, once again demonstrating that the reforming process is enthalpy-limited, meaning that the at the front portion of the catalyst the temperature is sufficiently high to drive steamreforming reactions, but the endothermic reactions decrease the catalyst temperature to the point where the catalyst is no longer active. The H2 concentration remaining nearly constant despite large transient changes in the other species indicates that the H2 concentration alone is not a complete descriptor of the reforming process.


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Figure 10. Transient response of the exit concentration of catalytic reforming products and monolith temperature profile at an O2 catalyst flow rate of 12 g/min and Φcatalyst = 7 (condition II in Figure 3). The transient response at condition III in Figure 3, shown in Figure 11, is at an O2 flow rate of 23 g/min and Φcatalyst = 10. As with condition II, there is a decrease in the mid-bed and outlet temperatures of more than 100°C that occurs within 2 min of the introduction of the postinjection. When the post-injection initially occurs, there is a spike in the CO, H2, and CH4 concentrations, followed by a rapid decline. Concurrently, there is an initial large decrease in H2O concentration, down to less than 2%, followed by an increase in its concentration. These concentration spikes then decrease as the catalyst outlet temperature decreases. It is believed that


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after the initial spike occurred, there was also an increase in the iso-octane concentration. However, the FTIR analyzer became saturated, and a reliable measurement of iso-octane at this condition was not possible with the measurement techniques used for this experiment.

Figure 11. Transient response of the exit concentration of catalytic reforming products and the monolith temperature profile at an O2 catalyst flow rate of 23 g/min and Φcatalyst = 10 (condition III in Figure 3). Thus, it is clear that the catalyst performance is dependent on the O2 flow rate and Φcatalyst condition. The performance can range from being exothermic at the lowest Φcatalyst investigated to being enthalpy-limited with a significant amount of unconverted fuel at the highest Φcatalyst. For the next section, the quasi-steady-state operation of the catalyst is compared at the various


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operating conditions. This considers only the data where the catalyst outlet temperature is within 10°C of the final temperature. Quasi-Steady-State Performance of Catalytic EGR-Loop Reforming Based on the equilibrium calculations described above, the major energy-containing reforming products are H2, CO, and CH4. The concentrations of each of these products are shown as functions of O2 flow rate and Φcatalyst in Figure 12 through 13. As illustrated in the previous section, steady-state reforming products are highly dependent on operating conditions, and Figure 12 shows that the catalyst-out H2 concentration ranges from 4% to 15%. The highest H2 concentrations occur for 4 < Φcatalyst < 7, and for O2 flow rates of 12 g/min and higher. For Φcatalyst > 7, steam reforming dominates but fuel conversion is enthalpy-limited, so overall H2 production is low. For Φcatalyst < 4, there is too much air and there is a shift to more complete oxidation, leading to water generation rather than H2 production. Figure 13 shows that the peak in CO concentration largely mirrors the H2 concentration, with the exception that the peak CO occurs at slightly leaner conditions: 3.5 < Φcatalyst < 5.5.


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Figure 12. H2 concentrations at the outlet of the catalyst as functions of Φcatalyst and O2 flow into the catalyst.

Figure 13. Carbon monoxide concentrations at the outlet of the catalyst as functions of Φcatalyst and O2 flow into the catalyst. Figure 14 reveals that the CH4 concentrations are low (about 300 ppm) for much of the experimental space investigated, but there is a high-magnitude band where catalyst-out CH4 was measured to be in-excess of 10,000 ppm. Initially this finding seems to be counter to the equilibrium modeling findings where CH4 was found to be highest at the highest O2 mole fractions and highest Φcatalyst. It is worth noting, though, that equilibrium calculations do not


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consider the timescales required to achieve equilibrium, nor do they consider whether the catalyst is an effective means of achieving equilibrium at all temperatures considered. There is only a short amount of time for reactions to occur in the catalyst (residence time < 100 ms, space velocity of 37,000 h-1), and as previously established in the Energetics and Equilibrium Analysis section, the high Φcatalyst and high O2 flow rate conditions are the most enthalpy-limited. If the endothermic reactions cause a sufficient temperature decrease, the catalyst activity will decrease and the CH4 concentrations will not reach equilibrium.

Figure 14. CH4 concentrations at the outlet of the catalyst as functions of Φcatalyst and O2 flow into the catalyst. Figure 15 shows the catalyst-out water concentration over the experimental space investigated. The water concentration contour plot is, in many ways, the inverse of that for H2 concentration in Figure 12. Instead of the maximum in H2 concentration, there is a minimum in water concentration for 4 < Φcatalyst < 7 and for O2 flow rates of 12 g/min and higher. Low water concentration is an indication that water is being consumed in the catalyst at some engine operating conditions. This is consistent with the water concentration observations in the previous section. For Figure 9, which was a condition dominated by POx reforming, there was very little


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change in the water concentration when the post-injection was introduced. However, for Figure 10 and Figure 11, which took place at higher Φcatalyst conditions, there was a large decrease in the catalyst-out water concentration.

Figure 15. H2O concentrations at the outlet of the catalyst with sweep of equivalence ratio in the catalyst and O2 flow rate into the catalyst. Water consumption can occur in two ways: steam reforming (eq 4) and the WGS reaction (eq 6). Further insight into the type of chemistry occurring over the catalyst can be gained by investigating the H2/CO ratio, shown in Figure 16. For iso-octane, the H2/COPOx = 1.125, and H2/COSteam = 2.125 from the stoichiometric ratios in Equations 3 and 4, respectively. Figure 16 shows that for all conditions investigated the H2/CO > H2/COPOx, and for Φ > 5.5, H2/CO > H2/COSteam. This is an indication that WGS chemistry does occur and plays an important role in increasing the overall H2/CO ratio at some conditions. Interestingly, the highest H2/CO ratios, which indicate more WGS chemistry, occur at the highest Φcatalyst conditions rather than where the highest H2 yields and the lowest H2O concentrations are observed.


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Figure 16. H2/CO at the outlet of the catalyst with sweep of equivalence ratio in the catalyst and O2 flow rate into the catalyst. Because the production of H2 in the WGS reaction in eq 6 is slightly exothermic, Le Chatelier’s principle suggests that a lower temperature will favor H2 production. Figure 17 shows the H2/CO ratio as a function of catalyst-out temperature. Interestingly, all O2 flow rate conditions collapse to show a very similar trend: when the exhaust temperature is below about 475°C, the H2/CO ratio is higher than what would be expected from steam reforming alone. This is consistent with the equilibrium calculations of H2/CO ratio shown in Figure 6. Thus, while WGS chemistry is present and it does shift the H2/CO ratio, it plays a greater role at low catalyst temperatures and high Φcatalyst, where the reforming activity is low.


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Figure 17. Correlation of H2/CO ratio and the catalyst-out temperature with sweep of equivalence ratio in the catalyst and O2 flow rate into the catalyst. (POx = partial oxidation; WGS = water-gas shift.) Thermal Analysis of the Catalyst Figure 9 through 11 illustrated that the catalyst mid-bed temperatures could either increase or decrease, depending on the dominant chemistry. Figure 18(a) and (b) show the temperature change across the first and second catalyst monoliths to provide a more direct indication of the type of chemistry that occurs under the conditions investigated. Over the first catalyst monolith, Figure 18(a), there is a temperature increase for Φcatalyst < 4 indicating exothermic chemistry, and there is a temperature decrease for Φcatalyst > 5 indicating endothermic chemistry. Without any reforming processes, there is normally a temperature rise of 10-20 °C over the first catalyst monolith due to oxidation of unburned species, as seen in Figure 9 through 11 prior to the postinjection. As a result, for the cases where a temperature decrease over the first catalyst monolith is measured, this can be attributed to the endothermic reforming processes. The second catalyst monolith, Figure 18(b), does not show any signs of exothermic chemistry. In contrast to the first catalyst monolith, the greatest temperature decreases are for the lowest Φcatalyst, and the highest Φcatalyst shows very little temperature change. This can again be attributed to enthalpy-limited chemistry. At a low Φcatalyst, the mid-bed temperature is sufficiently high to drive endothermic


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reactions in the second catalyst monolith. However, at a high Φcatalyst, all of the endothermic chemistry in the first catalyst monolith has already decreased the temperature significantly so that there is very little activity in the second catalyst monolith.

(a)

(b)

Figure 18. Temperature change in the first and second catalyst: (a) mid-bed temperature minus inlet temperature and (b) outlet temperature minus mid-bed temperature. Summary and Conclusions In this investigation, a combination of equilibrium calculations and fuel reforming on an engine were used to characterize catalytic reforming behavior and energetics over a range of Φcatalyst and O2 concentration conditions. Calculations of the reaction energetics provided insight into the desirable conditions from a thermodynamic standpoint. From an ideal standpoint, the highest levels of TCR are possible for the highest catalyst equivalence ratio. However, steam reforming is highly endothermic, and the system becomes enthalpy-limited, resulting in the production of CH4 as a reforming product rather than H2 and CO, which are more favorable from an energetics standpoint.


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Experimental results illustrated that high concentrations of H2 and CO could be formed over the catalyst at quasi-steady-state conditions, with as much as 15 vol % H2 at the catalyst outlet. There was a range of Φcatalyst and O2 flow rate conditions with which to achieve maximum reforming performance. The best performance was found for 4 < Φcatalyst < 7. Under these conditions the H2 and CO yield was highest, and measurements showed a significant amount of water consumption over the catalyst, confirming the presence of steam reforming reactions. Thermal analyses of the catalyst indicate that the reactions can be either endothermic or exothermic and also show that the majority of the reactions occur in the first catalyst monolith, indicating that higher space velocities are possible. In a companion article, the effect of this reforming strategy on the full multicylinder engine is reported.49 Corresponding Author *James Szybist; [email protected] Author Contributions The manuscript was written through contributions of all authors. All authors have given approval to the final version of the manuscript. ‡These authors contributed equally. (match statement to author names with a symbol) DISCLAIMER This manuscript has been authored by UT-Battelle, LLC, under Contract No. DEAC0500OR22725 with the US Department of Energy (DOE). The US government retains and the publisher, by accepting the article for publication, acknowledges that the US government retains a nonexclusive, paid-up, irrevocable, worldwide license to publish or reproduce the published form of this manuscript, or allow others to do so, for US government purposes. DOE


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will provide public access to these results of federally sponsored research in accordance with the DOE Public Access Plan (http://energy.gov/downloads/ doe-public-access-plan). ACKNOWLEDGMENT The authors gratefully acknowledge the support of the US Department of Energy Vehicle Technologies Office, particularly program managers Gurpreet Singh and Mike Weismiller. The authors also gratefully acknowledge John Nunan of Umicore for providing catalyst samples and Galen Fisher of the University of Michigan for supporting our efforts to obtain substrates and catalysts.

Yan Chang was also enrolled as a PhD candidate at the University of Michigan at the time of this publication. She would like to express gratitude for the strong support she received from her coadvisors, Prof. Andre Boehman and Dr. Stani Bohac.

ABBREVIATIONS BMEP, brake mean effective pressure; CA, crank angle; D-EGR, Dedicated EGR; EGR, exhaust gas recirculation; FSN, filter smoke number; FTIR, Fourier transform infrared; POx, partial oxidation; TCR, thermochemical recuperation; TDC, top dead center; WGS, water-gas shift; Φcatalyst, equivalence ratio at the catalyst inlet; Φcombustion, equivalence ratio in the combustion chamber for the main combustion event REFERENCES 1. Splitter, D.; Pawlowski, A.; Wagner, R. A Historical Analysis of the Co-evolution of Gasoline Octane Number and Spark-Ignition Engines. Frontiers in Mechanical Engineering 2016, 1 (1). 2. Alger, T.; Chauvet, T.; Dimitrova, Z. Synergies between High EGR Operation and GDI


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Systems. SAE Int. J. Engines 2009, 1 (1), 101-114, doi:10.4271/2008-01-0134. 3. Caton, J. A. A Comparison of Lean Operation and Exhaust Gas Recirculation: Thermodynamic Reasons for the Increases of Efficiency. SAE Technical Paper 2013, 201301-0266, doi:10.4271/2013-01-0266. 4. Alger, T.; Mangold, B.; Roberts, C.; Gingrich, J. The interaction of fuel anti-knock index and cooled EGR on engine performance and efficiency. SAE Int. J. Engines 2012, 5 (3), 12291241, doi:10.4271/2012-01-1149. 5. Tabata, M.; Yamamoto, T.; Fukube, T. Improving NOx and Fuel Economy for Mixture Injected SI Engine with EGR. SAE Technical Paper 1995, 950684, doi:10.4271/950684. 6. Kaul, B.; Finney, C.; Wagner, R.; Edwards, M.. Effects of External EGR Loop on Cycle-toCycle Dynamics of Dilute SI Combustion. SAE Int. J. Engines 2014, 7 (2), 606-614, doi:10.4271/2014-01-1236. 7. Cairns, A.; Blaxill, H.; Irlam, G. Exhaust gas recirculation for improved part and full load fuel economy in a turbocharged gasoline engine. SAE Technical Paper 2006, 2006-01-0047, doi:10.4271/2006-01-0047. 8. Duchaussoy, Y.; Lefebvre, A.; Bonetto, R. Dilution Interest on Turbocharged SI Engine Combustion. SAE Technical Paper 2003, 2003-01-0629, doi:10.4271/2003-01-0629. 9. Martz, J. B.; Middleton, R. J.; Lavoie, G. A.; Babajimopoulos, A.; Assanis, D. N. A computational study and correlation of premixed isooctane–air laminar reaction front properties under spark ignited and spark assisted compression ignition engine conditions. Combustion and Flame 2011, 158 (6), 1089-1096. 10. Stone, C. R.; Brown, A. G.; Beckwith, P. Cycle-by-cycle variations in spark ignition engine combustion-part ii: Modelling of flame kernel displacements as a cause of cycle-by-cycle variations. SAE Technical Paper 1996, 960613, doi:10.4271/960613. 11. Szybist, J. P.; Splitter, D. Effects of fuel composition on EGR dilution tolerance in spark ignited engines. SAE Int. J. Engines 2016, 9 (2), 819-831, doi:10.4271/2016-01-0715. 12. Alger, T.; Gingrich, J.; Mangold, B.; Roberts, C. A Continuous Discharge Ignition System for EGR Limit Extension in SI Engines. SAE Int. J. Engines 2011, 4 (1), 677−692, doi: 10.4271/2011-01-0661. 13. Wheeler, J.; Polovina, D.; Ramanathan, S.; Roth, K. Increasing EGR Tolerance using High Tumble in a Modern GTDI Engine for Improved Low-Speed Performance. SAE Technical Paper 2013, 2013-01-1123, doi:10.4271/2013-01-1123. 14. Conte, E.; Boulouchos K. Hydrogen-enhanced gasoline stratified combustion in SI-DI engines. Journal of Engineering for Gas Turbines and Power-Transactions 2008, 130 (2).


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15. Ji, C.W.; Wang, S.F. Effect of hydrogen addition on combustion and emissions performance of a spark ignition gasoline engine at lean conditions. International Journal of Hydrogen Energy 2009, 34 (18), 7823-7834. 16. D'Andrea, T.; Henshaw P.; Ting D.S.K. The addition of hydrogen to a gasoline-fuelled SI engine. International Journal of Hydrogen Energy 2004, 29 (14), 1541-1552. 17. Tahtouh, T.; Halter, F.; Samson, E.; Mounaïm-Rousselle, C. Effects of hydrogen addition under lean and diluted conditions on combustion characteristics and emissions in a sparkignition engine. International Journal of Engine Research 2011, 12 (5), 466-483. 18. Jamal, Y.; Wyszynski, M. L. On-Board Generation of Hydrogen-Rich Gaseous Fuels - A Review. Int. J. Hydrogen Energy 1994, 19, 557-572. 19. Fennell, D.; Herreros, J.; Tsolakis, A. Improving Gasoline Direct Injection (GDI) Engine Efficiency and Emissions with Hydrogen from Exhaust Gas Fuel Reforming. Int. J. Hydrogen Energy 2014, 39, 5153-5162. 20. Allenby, S.; Chang, W.-C.; Wyszynski, M. L. Hydrogen Enrighment: A way to Maintain Combustion Stability in a Natural Gas Fuelled Engine with Exhaust Gas Recirculation, the Potential of Fuel Reforming. Proc Instn Mech Engrs Part D 2001, 215, 405-418. 21. Alger, T.; Gingrich, J.; Mangold, B. The Effect of Hydrogen Enrichment on EGR Tolerance in Spark Ignited Engines. SAE Technical Paper 2007, 2007-01-0475, doi:10.4271/2007-010475. 22. Fennell, D.; Herreros, M.; Tsolakis, A.; Xu, H.; Cockle, K.; Millington, P. GDI Engine Performance and Emissions with Reformed Exhaust Gas Recirculation (REGR). SAE Technical Paper 2013-01-0537 2013. 23. Ivanič, Ž.; Ayala, F.; Goldwitz, J.; Heywood, J. Effects of Hydrogen Enhancement on Efficiency and NOx Emissions of Lean and EGR-Diluted Mixtures in a SI Engine. SAE Technical Paper 2005, 2005-01-0253, doi:10.4271/2005-01-0253. 24. Alger, T.; Mangold, B. Dedicated EGR: A New Concept in High Efficiency Engines. SAE Int. J. Engines 2009, 2 (1), 620-631. 25. Alger, T.; Walls, M.; Chadwell, C.; Joo, S.; Denton, B.; Kleinow, K.; Robertson, D. The Interaction bewteen Fuel AntiKnock Index and Reformation Ratio in an Engine Equipped with Dedicated EGR. SAE Int. J. Engines 2016, 9 (2). 26. Robertson, D.; Chadwell, C.; Alger, T.; Zuehl, J.; Gukelberger, R.; Denton, B.; Smith, I. Dedicated EGR Vehicle Demonstration. SAE Int. J. Engines 2017, 10 (3). 27. Chang, Y.; Wooldridge, M.; Bohac, S. Extending the Dilution Limit of Spark Ignition Combustion via Fuel Injection during Negative Valve Overlap. SAE Technical Paper 2016,


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2016-01-0671, doi:10.4271/2016-01-0671. 28. Szybist, J.; Steeper, R.; Splitter, D.; Kalaskar, V.; Pihll J.; Daw C. Negative Valve Overlap Reforming Chemistry in Low-Oxygen Environments. SAE Int. J. Engines 2014, 7 (1), 418433, doi:10.4271/2014-01-1188. 29. Fitzgerald, R.; Steeper, R. Thermal and Chemical Effects of NVO Fuel Injection on HCCI Combustion. SAE Int. J. Engines 2010, 3 (1), 46-64. 30. Urushihara, T.; Hiraya, K.; Kakuhou, A.; Itoh, T. Expansion of HCCI Operationg Region by the Combination of Direct Fuel Injection, Negative Valve Overlap, and Internal Fuel Reformation. SAE Technical Paper 2003-01-0749 2003. 31. Berntsson, A.; Denbratt, I. Optical Study of HCCI Combustion using NVO and an SI Stratified Charge. SAE Technical Paper 2007-24-0012 2007. 32. Ekoto, I.; Peterson, B.; Szybist, J.; Northrop, W. Analysis of Thermal and Chemical Effects on Negative Valve Overlap Period Energy Recovery for Low-Temperature Gasoline Combustion. SAE Int. J. Engines 2015, 8 (5), 2227-2239, doi:10.4271/2015-24-2451. 33. Ashida, K.; Maeda, H.; Araki, T.; Hoshino, M.; Hiraya, K.; Izumi, T.; Yasuoka, M. Study of aOn-board Fuel Reformer and Hydrogen-Added EGR Combustion in a Gasoline Engine. SAE Int. J. Fuels Lubr. 2015, 8 (2), 358-366, doi:10.4271/2015-01-0902. 34. Wheeler, J. C.; Stein, R.; Morgenstern, D.; Sall, E.; Taylor, J. Low-Temperature Ethanol Reforming: A Multi-Cylinder Engine Demonstration. SAE Technical Paper 2011-01-0142 2011. 35. Fowler, J.; Morgenstern, D.; Sall, E.; Veinbergs, M. Integration of an E85 Reforming System into a Vehicle-Ready Package and Project Results. SAE Technical Paper 2014, 2014-011191, doi:10.4271/2014-01-1191. 36. Hwang, J.; Li, X.; W., N. Exploration of Dual Fuel Diesel Engine Operation with On-Board Fuel Reforming. SAE Technical Paper 2017-01-0757 2017. 37. Fennell, D.; Herreros Arellano, J. M.; Tsolakis, A.; Wyszynski, M.; Cockle, K.; Pignon, J.; Millington, P. On-board Thermochemical Energy Recovery Technology for Low Carbon Clean gasoline Direct Injection Engine Powered Vehicles. Proc. IMechE Part D: J Automobile Enginnering 2017, 1-13. 38. Leung, P.; Tsolakis, A.; Rodriguez-Fernandez, J.; Goluski, S. Raising teh Fuel Heating Value and Recovering Exhaust Heat by On-Board Oxidative Reforming of Bioethanol. Energy Environ. Sci. 2010, 3, 780-788. 39. Jamal, Y.; Wagner, T.; Wyszynski, M. L. Exhaust Gas Reforming of Gasoline at Moderate Temperatures. Int. J. Hydrogen Energy 1996, 21 (6), 507-519.


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40. Gomes, S. R.; Bion, N.; Blanchard, G.; Rousseau, S.; Belliere-Baca, V.; Duprez, D.; Epron, F. Thermodynamic and Experimental Studies of Catalytic Reforming of Exhaust Gas Recirculation in Gasoline Engines. Applied Catalysis B: Environmental 2011, 102, 44-53. 41. Bogarra, M.; Herreros, J. M.; Tsolakis, A.; York, A. P. E.; Millington, P. J.; Martos, F. J. Impact of Exhaust Gas Fuel Reforming and Exhaust Gas Recirculation on Particulate Matter Morphology in Gasoline Direct Injection Engine. J. Aerosol Science 2017, 103, 1-14. 42. Bogarra, M.; Herreros, J. M.; Tsolakis, A.; York, A. P. E.; Millington, P. J.; Martos, F. J. Influence of On-Board Produced Hydrogen and Three Way Catalyst on Soot Nanostructure in Gasoline Direct Injection Engines. Carbon 2017, 120, 326-336. 43. Golunski, S. What is the Point of On-Board Fuel Reforming? Energy Environ. Sci. 2010, 3, 1918-1923. 44. Ahmed, S.; Krumpelt, M. Hydrogen from Hydrocarbon Fuels for Fuel Cells. Int. J. Hydrogen Energy 2001, 26, 291-301. 45. Chakravarthy, V. K.; Daw, C. S.; Pihl, J. A.; Conklin, J. C. Study of the Theoretical Potential of Thermochemical Exhaust Heat Recuperation for Internal Combustion Engines. Energy & Fuels 2010, 24, 1529-1537. 46. Tsolakis, A.; Megaritis, A.; Wyszynski, M. L. Application of Exhaust Gas Fuel Reforming in Compression Ignition Engines Fueled by Diesel and Biodiesel Fuel Mixtures. Energy & Fuels 2003, 17 (6), 1464-1473. 47. Tsolakis, A.; Megaritis, A.; Golunski, S. E. Reaction Profiles During Exhaust-Assisted Reforming of Diesel Engine Fuels. Energy & Fuels 2005, 19 (3), 744-752. 48. Brookshear, W.; Pihl, J.; Szybist, J. Catalytic Steam and Partial Oxidation Reforming of Liquid Fuels for Application in Improving the Efficiency of Internal Combustion Engines. Submitted for Publication Consideration to Energy & Fuels, August 2017. 49. Chang, Y.; Szybist, J.; Pihl, J.; Brookshear, D. Catalytic EGR-Loop Reforming for High Efficiency in a Stoichiometric SI Engine through TCR and Dilution Limit Extension 2. Engine Performance. Submitted for Publication Consideration to Energy & Fuels, August 2017. 50. Mehl, M.; Pitz, W. J.; Westbrook, C. K.; Curran, H. J. Kinetic modeling of gasoline surrogate components and mixtures under engine conditions. Proc. Comb. Inst. 2011, 33 (1). 51. Szybist, J. P.; Chakravathy, K.; Daw, C. S. Analysis of the Impact of Selected Fuel Thermochemical Properties on Internal Combustion Engine Efficiency. Energy & Fuels 2012, 26, 2798-2810.


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Catalytic Exhaust Gas Recirculation-Loop Reforming for High

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