Numerical Modeling and Simulation of Oxy-Combustion Exhaust Gas

explore the advantage of using the gas turbine high exhaust temperature and utilize the unwanted. CO2 for fuel upgrade via methane-exhaust gas reformi...
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Numerical Modeling and Simulation of Oxy-Combustion Exhaust Gas Recycling for Fuel Reforming Yinka S. Sanusi and Esmail M. A. Mokheimer* Mechanical Engineering Department, College of Engineering, KFUPM, Dhahran 31261, Saudi Arabia ABSTRACT: Different pragmatic approaches have recently been adopted in Paris Agreement of 2015 on the reduction of greenhouse gases, especially carbon dioxide (CO2). A viable option toward reduction of CO2 emission is to couple an exhaust gas fuel reforming reactor to an oxy-combustion power plant for in situ recycling and utilization of unwanted CO2 emission for fuel upgrade. In such a system, steam and dry fuel reforming take place simultaneously with the exhaust gases that are mainly water vapor and carbon dioxide. This study was carried out to explore the advantage of using the gas turbine high exhaust temperature and utilize the unwanted CO2 for fuel upgrade via methane-exhaust gas reforming. A numerical model was developed to study the effect of reforming temperature, gas hourly space velocity (GHSV) and reformer gas ratio (CO2/H2O) on the following performance metrics: methane conversion, fuel upgrade, and hydrogen yield in an exhaust gas-reforming reactor. Results show that the methane conversion increases with decreasing GHSV (i.e., increasing residence time) and approaches its thermodynamic equilibrium at GHSV= 1000 h−1. Different scenarios were investigated under different reformer gas ratio (CO2/H2O). At low temperature, steam methane reforming is the dominant route for methane conversion. While at high reforming temperature, methane becomes the limiting reactant that is consumed via both steam and dry reforming. Wet exhaust gas reforming with reformer gas ratio (CO2/H2O) of 2.85 has fuel upgrade of about 25% as compared to about 19% obtained in dry exhaust gas reforming. It was further shown that the reformer gas ratio (CO2/H2O) can be manipulated to give the desired syngas ratio (H2/ CO) depending on the application. For reformer gas ratio of 2.85, the syngas ratio obtained is 1.66 at 1073 K, which is suitable for use in Fischer−Tropsch and MeOH/DME synthesis. and oxygenated compound.8 The syngas produced via DMR can also be used as a flame stabilizer in low temperature gas turbines.9 Researchers have recently focused on combining SMR and DMR with emphasis on the development of catalyst that will give maximum methane conversion with negligible carbon deposit.10−13 Nickel-based catalyst that resists coke deposition14,15 has been tested successfully under combined steam and dry methane reforming (CSDMR). Jabbour16 showed that nickel−alumina catalyst (5 wt% Ni) synthesized using the evaporation-induced self-assembly method are more stable in combined steam and dry reforming of methane than those produced from conventional impregnated techniques. Ryi17 has also used catalytic nickel membrane to study CSDMR for application in gas to liquid process. They reported that no carbon deposit was observed on the nickel membrane. Saudi Arabia is blessed with abundant solar energy with an average annual solar radiation of about 2200 kWh/m2.18 Several solar researches including solar reforming have been commissioned by the Saudi government to explore the use of the abundant solar energy in the area of power and energy storage while, decreasing CO2 emission. The main challenge in solar reforming is the high temperatures required for the reforming reaction (>650 °C).19 This makes only point focusing, such as solar power tower, most suitable to provide the necessary energy. Rubin et al.20 recently carried out carbon dioxide reforming in solar tower of Weizmann Institute of Science.

1. INTRODUCTION About 32.19 Giga tones of CO2 were emitted via fuel combustion in 2013, representing about 59.1% increase in the years 1990−2013, and contributing more than 90% of the global anthropogenic GHG emissions.1 Several post Kyoto strategies in reducing CO2 emission from power plants are categorized as postcombustion, precombustion, and oxy-fuel combustion techniques.2 The oxy-fuel combustion is reported to be a more promising carbon capture technique,3 when the cost and energy penalties in CO2 capture are considered. The exhaust products from oxy-fuel power plant are mostly carbon dioxide and water vapor. These exhaust products can be used in upgrading hydrocarbon fuel via reforming. Moreover, the exhaust gases from commercial gas turbines are usually at high temperature and can be more than 500 °C.4,5 The gas turbine exhaust temperature could be higher under oxy-fuel combustion. The direct use of exhaust gas in fuel reforming will enable the heat recovery from the exhaust gases and reduce the emission of unwanted CO2 to the atmosphere. Figure 1 shows, conceptually, the integration of a reformer bed to oxy-fuel gas turbine plant. The exhaust that is predominantly water vapor and CO2 can be used to reform natural gas fuel to produce syngas fuel. Steam and dry reforming of methane (which is main constituent of natural gas) has been widely investigated due to abundance of natural gas. In steam methane reforming (SMR), water vapor reacts with methane in the presence of catalyst to produce hydrogen rich syngas.6,7 Unlike the SMR, the dry methane reforming (DMR) uses the unwanted greenhouse gas (CO2) for the production of syngas of lower H2/CO ratio than that in SMR that is suitable for the production of liquid hydrocarbon (Fischer−Tropsch synthesis) © 2017 American Chemical Society

Received: March 16, 2017 Revised: April 13, 2017 Published: April 13, 2017 5385

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Figure 1. Conceptual system integration of oxy-natural gas plant and reformer bed.

flame temperatures. The percentage of recycled exhaust gases will vary depending on combustor design and material limitation. This leads to varying exhaust gas temperature and compositions. The effect of exhaust gas composition at different reformer operating conditions, such as temperature, feed rate etc., needs to be well understood ahead of system integration. Consequently, this study was carried out to investigate combined steam and dry methane reforming (CSDMR) numerically in a tubular packed bed reactor under industrial relevant conditions (reformer gas composition, temperature, feed rate), using computational fluid dynamics approach. A steady state model was developed and was first used to compare the performance of two reactors: the heated reactor and preheated gas reactor at atmospheric conditions. In preheated gas reactors, the reforming gases (exhaust gases + methane) are heated prior to entering the adiabatic reformer bed. On the other hand, heat is directly supplied to the reformer bed via its wall in the heated reactor. The effect of varying operating conditions, such as operating temperature between 773 to 1073 K, gas hourly space velocity (GHSV) between 300 h−1 to 10 000 h−1 and different cases of oxycombustion scenarios were also considered. The gas composition at the inlet is adjusted such that reformer gas (H2O + CO2) to methane ratio is 3 on molar basis. While the reformer gas ratio (CO2 /H2O) varies between 0.5 and 2.85 depending on oxy-combustion scenarios considered. The GHSV was computed as given in eq 1, while the metric used in assessing the performance of the reactor under the listed operating conditions are the methane conversion (XCH4),

They achieved absorber temperature of about 1450 K and 85% methane conversion. Such solar tower can be integrated to the exhaust of oxy-combustion of gas turbine power plant for reforming of natural gas. It is imperative to state that aircombustion exhaust are not good candidate for exhaust gas reforming due to the presence of excess air that can aid undesirable combustion of natural in the reformer bed. Unlike air combustor, oxy-combustion plants are designed to operate at near stoichiometric condition due to high cost of oxygen production. The negligible oxygen concentration in their exhaust makes it a preferred candidate for exhaust gas reforming. 1.1. Integration of Reformer Bed to Oxy-Natural Gas Plant. The conceptual system integration for natural gas reformer bed to oxy-natural gas turbine power plant is shown in Figure 1. According to the figure, natural gas is burnt in an oxycombustor (i.e., combustion chamber) in the presence of oxygen and recycled exhaust gas. The combustion product is expanded in a gas turbine to generate power or used to generate steam in a boiler for power and other industrial purposes. The turbine/boiler exhaust which is still at high temperature is used directly for natural gas reforming to produce syngas. Alternatively, water vapor in the turbine/boiler exhaust is condensed such that pure natural gas dry reforming (CO2) is carried out in the reformer bed. The auxiliary heat energy required is sourced from solar energy leading to conversion of solar energy to chemical energy (produced syngas). The advantages of integrating reformer bed to oxycombustor gas turbine power plants over conventional carbon capture system include: (1) reducing the adverse environmental effect of CO2 emission to the atmosphere and converting it to a value-added product, (2) reducing the cost of transportation and of the sequestration of CO2, (3) enhancing heat recovery from the exhaust gases, (4) storing solar energy in a thermochemical form, and (5) producing syngas of desired H2/CO composition depending on application. This article is focused on investigating the performance of reformer bed in oxy-combustor exhaustnatural gas reforming. 1.2. Objectives of the Present Study. In practical oxyfuel combustion, part of the exhaust gases is recycled to lower

H ( CO ), hydrogen yield per mole of methane supplied

syngas ratio

2

(Hy), and percentage of fuel up grade (1 F) given in eqs 2, 3, 4, and 5, respectively. GHSV =

Q̇ Vreactor

Methane conversion (X CH4) = 5386

(1)

CH4,in − CH4,out CH4,in

(2)

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Energy & Fuels Syngas ratio (Sr) =

n

mole of H 2 produced mole of CO produced

hydrogen yield (Hy) =

mole of H 2 produced CH4,in

(3)

∇ . (ερg uY i i,j ̅ j) = ∇ . (ρg Dj . ∇Yj) + (1 − ε)ρc Mj ∑ rv

(4)

β = 150

Percentage of fuel up grade (1F) ⎛ (M . LHV + M . LHV + M · LHV ) ⎞ H2 H2 CO CO CH 4 CH 4 out = ⎜⎜ − 1⎟⎟ × 100 MCH4,in ·LHVCH4 ⎝ ⎠

(1 − ε)2 μg ρg ε 3d p2

+ 1.75

(1 − ε)|u ̅ | ε 2d p

(10)

2.3. Reaction Kinetics. The exhaust gas reforming investigated in this study is a combination of steam and dry methane reforming. The equations that govern this reforming process are Steam methane reforming reactions (SMR):29

(5)

2. NUMERICAL MODEL 2.1. Reactor Geometry. This article is focused on the use of the gas turbine/boiler oxy-combustion exhaust gases to reform natural gas. It is important to state that oxy-combustion exhaust gases are considered for exhaust gas reforming, since their combustors are operated at near stoichiometric condition due to high cost of oxygen production resulting in negligible oxygen in the exhaust. Recycled exhaust gases are, thus, used to maintain the combustor temperature and enhance turbulence in such combustor. For simplicity, this study was carried out for exhaust gas reforming of methane (main constituent of natural gas) in a tubular packed bed reactor. This kind of reactor has been widely studied for both steam21,22 and dry23,24 methane reforming. The packed bed reactor adopted in this study is shown in Figure 2. The reactor has a length of 200 mm with

ΔH o298K = 206 KJ/mol

CH4 + H 2O ↔ CO + 3H 2

(11)

CH4 + 2H 2O ↔ CO2 + 4H 2

ΔH o298K

= 164.9 KJ/mol (12)

Dry (CO2) methane reforming reaction (DMR):

28

ΔH o298K = 247.3 KJ/mol

CH4 + CO2 ↔ 2CO + 2H 2

(13)

Water−gas shift reaction (WR):29 CO + H 2O ↔ CO2 + H 2

ΔH o298K = − 41.1 KJ/mol (14)

The WR is a very important reaction that occurs in both the SMR and DMR. This reaction (WR) has a significant impact on the overall reforming process. The reaction kinetics developed by Xu and Forment30 for steam methane reforming reactions (eqs 11, 12, 14) over nickel-supported catalyst are given in eqs 15, 16, and 18, respectively. These equations have been extensively used in SMR modeling in both laboratory and industrial scales.28,29,31,32 In modeling dry reforming of methane, we considered the reaction kinetics (eq 17) that were initially developed for Rh/Al2O3 catalyst by Richardson and Paripatyadar.33 Benguerba9 has shown earlier that the reaction kinetics developed by Richardson and Paripatyadar33 predicted their experimental data with Ni/Al2O3 catalyst accurately. Our validation, presented in Section 3, also shows that the dry reforming reaction kinetics33 is applicable to a nickel-base catalyst.

Figure 2. Packed bed reactor.

reactor length to diameter aspect ratio of 10 and reactor diameter to particle diameter ratio of ∼7. A tube to particle diameter ratio between 4 and 15 is recommended due to pressure drop constraints.25 The adopted reactor is similar to the one reported in reference.24 A computational fluid dynamics approach was used to model the reactor under steady state conditions. The reactor is modeled as a pseudohomogeneous catalytic bed such that the feed gases and the catalyst are modeled as one continuum. Details of the mathematical equations used to model the reactor are described in Sections 2.2 and 2.3. 2.2. Mathematical Model. The reactor is modeled using the conservation equations of mass (eq 6), momentum (eq 7), energy (eq 8), and species (eq 9).26,27 The pressure drop in the reactor due to local friction was computed using the Ergun equation28 as given in eq 10. ∇ . ερg u ⃗ = 0

r1 =

r2 =

k1 PH2.52 k2 PH3.52

⎛ PCOPH32 ⎞ 1 ⎜ ⎟ − P P CH H O Keq,1 ⎟⎠ DEN2 ⎜⎝ 4 2

(15)

⎛ PCO2PH42 ⎞ 1 ⎜ 2 ⎟ PCH PH O − Keq,2 ⎟⎠ DEN2 ⎜⎝ 4 2

(16)

⎛ ⎞⎛ 2 ⎞ KD,CO2KD,CH4PCO2PCH4 ⎟⎜1 − (PCOPH2) ⎟ r3 = k 3⎜⎜ 2 ⎟⎜ Keq,3PCO2PCH4 ⎟⎠ ⎝ (1 + KD,CO2PCO2 + KD,CH4PCH4) ⎠⎝

r4 =

PCO2PH2 ⎞ k4 1 ⎛ ⎟ ⎜ P P − CO H O ⎜ 2 Keq,4 ⎟⎠ PH2 DEN2 ⎝

(6)

∇ . ερg uu⃗ ⃗ = −ε∇P − βερg u ⃗ + ∇ . ετ ⃗

(9)

i=1

DEN = 1 + K CH4PCH4 + K COPCO + K H2PH2 +

(7)

(17)

(18)

PH2OK H2O PCH4 (19)

n

c p,g∇ . (ερg uT ̅ ) = ∇ . (K . ∇T ) + (1 − ε)ρc ∑ riΔHi

The reaction rate constant ki, equilibrium constant Keq,i, and absorption constant Kj are computed based on Arrhenius equation as given below:

i=1

(8) 5387

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mm long and it is packed with nickel catalyst supported on gamma alumina. They reported that the residence time in all

(20)

their experiment is 3.6

The Arrhenius parameters (i.e., the pre-exponential factor and the activation energy) for the reaction rate constant ki, equilibrium constant Keq,i, and absorption constant Kj are summarized in Tables 1, 2, and 3, respectively.

kgcat·s mol CH4

. They did not observe carbon

deposition on the catalyst signifying that catalyst deactivation can be ignored as done in the present modeling. In Figure 3, we

Table 1. Arrhenius Parameters for Reaction Rate

(

E

(

Constantki ki = Ai exp − R·Ti

))

i

Ai (dimension of ki)

Ei (KJ/mol)

source

1 2 3 4

−1 4.225 × 10 (Kmol bar kg−1 cat h ) −1 1.02 × 1015 (Kmol bar5 kg−1 h ) cat −1 1290 (mol g−1 cat s ) −1 1.955 × 106 (Kmol bar−1 kg−1 cat h )

240.1 243.9 102.065 67.13

30 30 33 30

15

5

Table 2. Arrhenius Parameters for Equilibrium Rate

(

( )) C

i

Bi (dimension of Bi)

1 2 3 4

1.198 2.117 6.780 1.767

× × × ×

Figure 3. Comparison of the experimental31 and predicted methane conversion (XCH4, red dot), wet concentration of hydrogen (YH2, light blue diamond), and carbon monoxide (YCO, dark blue square) at different temperatures. Lines and markers: numerical, Markers only experiments.

34

Constant Keq, i Keq, i = Bi exp − Ti

Ci (KJ/mol)

1017 (kPa2) 1015 (kPa2) 1018 (kPa2) 10−2 (−)

26830 22430 31230 −4400

showed that our current model gave excellent predictions of the experimentally observed values of the methane conversion and syngas components (H2 and CO) at different reforming temperatures. For the case of dry methane reforming, experimental results from Lu24 were used to validate our model. Their reformer bed is of 200 mm long and 26 mm inner diameter packed with commercial nickel alumina catalyst. Our predicted methane conversion was compared with experimental data at different temperatures as given in Figure 4. The

Table 3. Arrhenius Parameters for the Species Absorption Fj

(

( ))

Rate Constant Kj Kj = Gjexp − R·T j CH4 H20 CO H2 D,CO2 D,CH4

Gi (dimension of Gi) −4

−1

6.65 × 10 (bar ) 1.77 × 105 (−) 8.23 × 10−4 (bar−1) 6.12 × 10−9 (bar−1) 2640 (Pa−1) 2630 (Pa−1)

Fi (KJ/mol)

source

−38.28 88.68 −70.61 −82.90 37.641 40.684

30 30 30 30 33 33

2.4. Computational Scheme. For simplicity, a 2-D axisymmetric model was developed based on the geometry shown in Figure 2 to study the performance characteristics of methane-exhaust gas reformer. The set of equations described earlier are solved numerically under appropriate boundary conditions using ANSYS fluent 15 code. Details of the equations coupling and solution procedures are detailed in.27,35 SIMPLE algorithm was used for pressure−velocity coupling while second order upwind was used for the advection term in the momentum, energy and species equations for enhanced accuracy. The convergence criteria for the continuity, momentum is achieved when the residual is less than 10−6. While the residual for the energy and species equations is set to be less than 10−9.

Figure 4. Comparison of the experimental24 and predicted methane conversion at different temperatures for DMR.

experimental values were well predicted at higher temperatures with over predictions at lower temperatures. Benguerba9 have similarly observed an over prediction of methane conversion at temperature below 600 °C while using reaction kinetics by Richardson and Paripatyadar33 for nickel catalyst. The summary of our validation results shows that the kinetic model predicted the experimental data within 5% at high temperature and an average of 20% over prediction at low temperature. This indicated that the CFD and kinetic model presented can be further used to model methane steam and dry reforming or its combination with nickel-based catalyst with good accuracy at high reforming temperature.

3. MODEL VALIDATION Validation runs were carried out for MSR and MDR to validate the set of equations presented in Section 2. Results obtained from the computational model (i.e., continuity, momentum, energy, species, and reaction kinetics) of steam and dry methane reforming were validated against the experiential data obtained from Hoang31 and Lu,24 respectively. In the work of Hoang,31 the reformer bed is of 10 mm inner diameter and 150 5388

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4. RESULTS AND DISCUSSION 4.1. Effect of Temperature. The reforming reactions are highly endothermic reactions that require substantial amount of energy for the catalytic activity to take place. This study is carried out to take advantage of high exhaust gas temperature and to utilize the unwanted CO2 for fuel upgrade. The first part of this work is focused on the effect of reactor temperature and mode of applying heat energy on the reactor performance. In this regard, two different scenarios were considered. The first is to maintain the reactor wall at a given temperature and is referred to as “heated reactor”. The second is to preheat the gas to a specified temperature while keeping the reactor wall at adiabatic condition and is referred to as “preheated gas reactor”. In both cases, the exhaust from a gas turbine power plant or a boiler with a pure oxy-combustion of methane was utilized in the reforming such that methane and reformer gas (CO2 + H2O) were supplied to the reactor. The reformer gas ratio (CO2/H2O) is 0.5. As observed in Figure 5, a near isothermal

Figure 6. Methane conversion at different temperatures for GHSV = 1000 h−1 and CO2/H2O = 0.5.

reactor, which was observed in heated reactor. The observed methane conversion for heated reactor falls within the thermodynamic conversion for steam and dry reforming of methane. For instance, at 1073 K our observed methane conversion is 98.6%, which falls within 100% and 90.8% corresponding to thermodynamic conversion of steam and dry reforming of methane. Also at intermediate temperature of 973 K, our observed methane conversion is 82.125%, which similarly falls within steam methane conversion of 97.4% and 72.8% for dry reforming (at the same temperature). This implies that the presence of steam in the exhaust gas results in higher methane conversion than what would have been achieved, if pure DMR of exhaust gas (CO2) has been considered. The contribution of the SMR to the overall exhaust gas reforming (EGR) is due to its lower energy requirement as compared to DMR per mole of CH4 reformed. The produced syngas (H2 and CO) has higher enthalpy than the methane supplied to the reformer. This implies that heat energy has been stored as chemical energy in the fuel. We observed similar trend in the fuel upgrade (see Figure 7) as earlier observed in the

Figure 5. Methane conversion and temperature distributions along the reactor center line.

condition was observed for large part of the heated reactor, while significant temperature drop across the reactor was observed for preheated gas reactor. The effect of the temperature distribution within the reactors is evident in the progress of methane conversion for both reactors. The dependence of methane conversion on temperature has been well documented in literatures17,24,29,36 and is also exhibited in the Arrhenius equation presented in Section 2.3. The large temperature drop in the case of the preheated gas reactor and continuous temperature drop downstream of the reactor result in slow progress in the methane conversion until the reactor exit. The low heat capacity of methane and reformer gases limit the amount of heat supplied to the preheated gas reactor far less than that required for 100% conversion even at 1073 K. The computed heat energy supplied via the wall of heated reactor is more than 4 times that provided by the feed gases in the preheated gas reactor at the same temperature (1073 K). Thus, methane conversion at the reactor exit for heated reactor is 98.6% which is far more than 32.2% obtainable in preheated gas reactor. Besides the huge difference in their methane conversion, preheated gas reactor that is characterized by large temperature gradient also has higher tendency for thermal stress in the reactor. These stresses are of great concerns in reforming reactors.37 The methane conversion is generally observed to increase with increasing temperature for both reactors as shows in Figure 6. The difference in the methane conversion for the two reactors increases with increasing temperature. This further justified the importance of near isothermal condition in the

Figure 7. Fuel upgrade at different temperatures for GHSV = 1000 h−1 and CO2/H2O = 0.5.

methane conversion for the two reactors. This is due to the fact that the fuel upgrade is directly related to the amount of gas reformed such that higher fuel upgrades are similarly observed at higher temperatures. At 1073 K, the fuel upgrade is about 21% for heated reactor as compared to about 5% observed in preheated gas reactor. In Figure 8, the hydrogen yield was also observed to increases with increasing the temperature in a similar fashion as that of methane conversion. This is due to the increase hydrogen selectivity with increasing temperature. For heated reactor at 1073 K, hydrogen yield is approximately 3 mol per mole of methane, which is the thermodynamic limit for steam methane reforming. This suggests that for the case where 5389

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of CH4 in the reactor, such that higher GHSV signifies lower residence time of CH4 in the reactor or low contact time between the feed gas and the catalyst. In Figures 10 and 11, we

Figure 8. Hydrogen yield at different temperatures for GHSV = 1000 h−1 and CO2/H2O = 0.5.

reformer gas ratio (CO2/H2O) is 0.5. The steam methane reforming is dominant route for the EGR. Thus, reforming of exhaust gas of this type can be treated as purely SMR and can be used for hydrogen production. The relative composition of the syngas produced (i.e., syngas ratio Sr) is presented in Figure 9. We observed higher syngas

Figure 10. Methane conversion and fuel upgrade at different GHSV. T = 1073 K and CO2/H2O = 0.5.

Figure 11. Hydrogen yield and syngas ratio at different GHSV. T = 1073 K and CO2/H2O = 0.52. Figure 9. Syngas ratio (H2/CO) at different temperatures for GHSV = 1000 h−1 and CO2/H2O = 0.5.

presented the effect of reactor GHSV on the performance of the heated reactor. The methane conversion increases with decreasing GHSV until GHSV of 1000 h−1, where further decrease in the GHSV only lead to a marginal increase in the methane conversion (see Figure 10). This indicated that the reactor approaches thermodynamic equilibrium at GHSV= 1000 h−1. Serrano-Lotina36 and Guo40 have equally reported that low methane conversion are observed at low contact time (high GHSV) between the feed gases and the catalyst which approaches equilibrium conversion at a critical contact time. In the present study we observed that by doubling the contact time (i.e., decrease GHSV from 1000 h−1 from 500 h−1), the methane conversion marginally increase by ∼0.9%. SerranoLotina36 reported that methane conversion are strongly dependent on the kinetics at low contact time (high GHSV) until equilibrium beyond which mass transfer dominates. This implies that at GHSV of more than 1000 h−1 and the operating conditions presented in Figure 10, the methane conversions are strongly influenced by the kinetics of H2 and CO selectivity by the Ni-based catalyst. The fuel upgrade and hydrogen yield were observed to follow similar trend for different GHSV as shown in Figures 10 and 11, respectively. At higher GHSV (say 10 000 h−1), the fuel upgrade and hydrogen yield are low and increases with decreasing GHSV up until optimum GHSV = 1000 h−1 is reached. Therefore, further decrease in the GHSV leads to decrease in both the fuel upgrade and hydrogen yield. The reduced fuel up grade at GHSV of less than 1000 h−1 could be attributed to the effect of inverse water gas shift reaction

ratio at lower temperature. This could be attributed to the poor CO selectivity at low temperatures. Wang38 has earlier reported that SMR which is dominant route for the methane reforming in the present case are characterized by poor CO selectivity. The poor CO selectivity is further compounded by the water gas shift reaction (eq 14) that is favored at low temperature leading to increase production of H2 and consumption of CO at lower temperature. Thus, higher syngas ratio were observed at low temperatures. Moreover, Soria39 demonstrated that at low temperatures (623−773 K) in a methane dry reforming with 5% H2O, water gas shift reaction is favored, leading to increasing syngas ratio (H2/CO) with decreasing temperature. We similarly observed that preheated gas reactor that are characterized by lower reformer temperature as compared to heated reactor has higher syngas ratio. At 1073 K, the syngas ratio in the heated reactor is about 3.63 which is more than thermodynamic limit of 3 for SMR indicating poor CO selectivity. Another reason that is associated with poor CO selectivity is the contact time between the feed gases and the catalyst, which is discussed in Section 4.4. It is, however, important to restate here that despite decreased syngas ratio (H2/CO) as the reformer temperature increases, the overall hydrogen produced increases, due to increased methane conversion. 4.2. Effect of Gas Hourly Space Velocity. Gas hourly space velocity (GHSV) was used to quantify the residence time 5390

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Energy & Fuels Table 4. Performance of the Reformer at Different Reformer Gas (CO2/H2O) Ratio methane conversion (%)

fuel upgrade (%)

temperature (K)

Case 1

Case 2

Case 3

Case 4

Case 1

Case 2

Case 3

Case 4

873 973 1073

49.09 82.13 98.58

48.15 82.73 99.26 hydrogen yield

46.92 82.32 99.57 (Hy)

43.92 78.97 99.81

10.07 17.73 21.04

10.03 9.89 18.23 18.50 21.77 22.72 syngas ratio (H2/CO)

9.39 18.38 24.38

temperature (K)

Case 1

Case 2

Case 3

Case 4

Case 1

Case 2

Case 3

Case 4

873 973 1073

1.56 2.55 3.00

1.50 2.51 2.92

1.44 2.42 2.76

1.29 2.16 2.41

4.51 3.88 3.64

4.11 3.50 3.13

3.74 3.08 2.49

3.19 2.37 1.66

reforming is 41%. This demonstrates that the increasing concentration of CO2 in the reformer gas slow down the kinetic conversion rate of CH4. However, at temperature of 1073 K, the methane conversion for all cases ranges between 98.58−99.81% with 99.81% achieved for Case 4 having higher CO2 concentration in the reformer gas. This indicates that dry reforming reaction is favored as it is similarly reported in.43 At high temperature, methane becomes the limiting reactant that is consumed via both steam and dry reforming. Ryi17 has also reported that at temperature ≥973 K, the methane conversion is independent of reformer gas ratio. Moreover, Choudhary44 stated that at high temperature of 1123 K the methane conversion is independent of reformer gas composition for reformer gas ratio of 0.05−1 but marginally increases to 99.3 from 98.5 as the reformer gas ratio is increased to 1.82. This is in conformity with our findings in which the methane conversion for all cases ranges between 98.5−99.8% for reformer gas ratio between 0.5 (for Case 1) to 2.85 (for Case 4) at 1073 K. Our calculations show that for 100% methane conversion, the maximum fuel grade expected for SMR and DMR are 26 and 32%, respectively. Thus, dry reforming of methane will result in higher fuel upgrade compared to steam reforming at the same methane conversion. For similar methane conversion, higher fuel upgrade is observed at higher reformer gas ratio (see Table 4 at 1073 K). This suggests that at high temperature both reformer gases (CO2 and H2O) are competing for methane consumption leading to increase methane dry reforming. The contribution of dry reforming to the overall reforming process leads to decrease hydrogen yield and syngas ratio (H2/CO) at higher reformer gas ratio as shown in Table 4. The syngas ratio at 1073 K decreases from ∼3.64 at reformer gas ratio of 0.5 to ∼1.66 at reformer gas ratio of 2.85, thus, approaching thermodynamic limit of 1 for DMR. This implies that reformer gas ratio can be manipulated to give the desired syngas ratio depending on the application. 4.4. Comparison of Reforming of Wet and Dry Exhaust Gas. The last scenario compared in this study is the reforming of dry and wet reforming of exhaust gas. The dry reforming can be carried out by condensing the water vapor in the oxy-combustion exhaust gas such that pure CO2 is used for reforming of CH4 (DMR). While the wet exhaust gas considered has reformer gas with 74% CO2 (26% H2O vapor) concentration (as Case 4). Methane conversion for Case 4 is generally more than those of dry reforming at all temperatures as shown in Figure 12. Higher methane conversion for case 4 is due to steam methane reforming contribution to the overall methane reforming. At 1073 K, the methane conversion for Case 4 is 99.8%, while that of dry

(WGSR) that favor the consumption of H2 in the presence of high volume of CO2 leading to the formation of CO and H2O. The effect of H2 consumption is evident in Figure 11, such that hydrogen yield decreases for GHSV < 1000 h−1 even with marginal increase in methane conversion. Unlike other performance metrics (i.e., methane conversion, fuel upgrade, and hydrogen yield), the CO selectivity increases monotonically (not shown here) with decreasing GHSV (increasing contact time). The increase of CO selectivity at lower GHSV resulted in decrease in syngas ratio (H2/CO), as given in Figure 11, despite the fact that hydrogen yield increases. At GHSV < 1000, we observed a drastic decrease in the syngas ratio produced. Purnama41 experimentally showed that there is distinct change in the CO formation rate at high contact time due to the change in the controlling kinetics of CO formation at methane conversion of more than 70%. They, however, concluded that the reverse water gas shift reaction can satisfactory account for CO selectivity at high contact time (low GHSV) despite the complex nature of the real reaction kinetics. The effect of the reverse water gas shift reaction (WGSR) that favors production and consumption of CO and H 2 , respectively, leads to sharp decrease in the syngas ratio (H2/ CO) at GHSV < 1000 h−1. 4.3. Effect of Reformer Gas (CO2/H2O) Ratio. In practical oxy-combustion, oxygen is diluted with recycled CO2 or exhaust gases (CO2/H2O) to lower the flame temperature. Thus, the reformer gas (CO2/H2O) ratio in the exhaust is higher compared to that of the case of pure-oxy combustion discussed in Section 4.1. In this section, we presented the results of different cases of methane reforming of exhaust gases. Case 1 is the base case presented in Section 4.1. It is based on pure oxy-combustion of methane (i.e., 0% CO2 dilution) that results in reformer gas (CO2/H2O) of ratio 0.5. Cases 2, 3, and 4 have reformer gas obtained from exhaust of oxy-combustion of methane with CO2 dilution of 30, 50, and 70% in the oxidizer resulting in reformer gas (CO2/H2O) ratio of 0.92, 1.5 and 2.85, respectively. Oxy-combustion of methane with 70% CO2 dilution has been previously reported to have similar combustion characteristics and adiabatic flame temperature with air combustion42 and its exhaust gas is referred to as Case 4. The performance of reactor under varying reformer gas ratio is summarized in Table 4. At low reforming temperature (873 K), the methane conversion decreases from ∼49% in Case 1 to ∼43% in Case 4. This is due to the increase concentration of CO2 in the reformer gas. Dry methane reforming is more endothermic than steam as illustrated in eqs 11 and 13. Thus, the catalytic conversion of CO2 reformer are lower than that of SMR at the same temperature. For instance at 873 K, the thermodynamic limit for SMR is 77% while that of dry 5391

DOI: 10.1021/acs.energyfuels.7b00772 Energy Fuels 2017, 31, 5385−5394

Article

Energy & Fuels

(DMR) reforming. We observed a marginal decrease in H2 yield for Case 4 beyond 1023 K. This is due to the fact that DMR becomes more dominant route for CH4 reforming than SMR that results in the production of less H2 even at higher methane conversion (see Figures 12 and 14). This suggests that the temperature of 1023 K could be the optimum temperature for hydrogen production. At all temperatures, the H2 yield is more than twice those of DMR. Despite the increase of hydrogen yield with temperature, the syngas ratio (H2/CO) monotonically decreases with temperature especially for Case 4, while the syngas increases marginally with temperature for methane dry reforming. Thermodynamically, the syngas ratio in a dry reforming is unity. We, however, observed values that are less than 1. Similar observations have been reported by previous researchers. The effect of reverse was water gas shift reaction that results in the consumption and production of H2 and CO, respectively, has been attributed to having syngas ratio