Analysis of Syngas Production from Biogas via the Tri-Reforming Process

Abstract

The tri-reforming process was employed for syngas production from biogas at elevated pressures in this study. In the tri-reforming process, air and water were added simultaneously as reactants in addition to the main biogas components. The effects of various operating parameters such as pressure, temperature and reactant composition on the reaction performance were studied numerically. From the simulated results, it was found that methane and carbon dioxide conversions can be enhanced and a higher hydrogen/carbon monoxide ratio can be obtained by increasing the amount of air. However, a decreased hydrogen yield could result due to the reverse water–gas shift reaction. A higher level of methane conversion and hydrogen/carbon monoxide ratio can be obtained with increased water addition. However, negative carbon dioxide conversion could result due to the water–gas shift and reverse carbon dioxide methanation reactions. The dry reforming reaction resulting in positive carbon dioxide conversion can only be found at a high reaction temperature. For all cases studied, low or negative carbon dioxide conversion was found because of carbon dioxide production from methane oxidation, water–gas shift, and reverse carbon dioxide methanation reactions. It was found that carbon dioxide conversion can be enhanced in the tri-reforming process by a small amount of added water. It was also found that first-law efficiency increased with increased reaction temperature because of higher hydrogen and carbon monoxide yields. Second-law efficiency was found to decrease with increased temperature because of higher exergy destruction due to a more complete chemical reaction at high temperatures.

1. Introduction

The efficient production of syngas (a mixture of hydrogen and carbon monoxide) is gaining significant attention worldwide as it is a versatile feedstock that can be used to produce a variety of fuels and chemicals, such as methanol, Fischer–Tropsch fuels, H₂, and dimethyl ether (DME) [1]. Using CH₄ as the primary material, syngas can be produced from steam reforming (SR), partial oxidation (POX), autothermal reforming (ATR), and dry reforming (DR). The tri-reforming (TR) process for syngas production from CH₄ has received growing attention because of its technical simplicity and flexible operation [2,3,4,5]. In the TR process, the syngas is produced by combining SR, DR, and POX in a single step. The TR process was proposed originally for syngas production from power plant flue gas [6,7]. There are several advantages for syngas production from the TR process. As CO₂ is one of the reactants, there is no need for CO₂ separation from the flue gas [6,7]. The H₂/CO ratio in syngas can be altered by adjusting the relative amounts of the reactants. In addition, the presence of H₂O and O₂ in the feedstock helps to mitigate carbon deposition, and catalyst deactivation can be prevented [8,9].

As fossil energy resources reduce sharply and environmental pollution becomes more serious, searching for new materials for syngas production plays an important role in future energy development [10,11,12]. Biogas is receiving much attention because of its considerable economic and environmental benefits [13]. The biogas composition is related to the starting substrate, but is basically composed of CH₄ and CO₂ with the volume ratio of 2 [14,15,16]. Both CH₄ and CO₂ are regarded as major greenhouse gases (GHGs), which pose a serious threat to the global climate and environment. Using biogas for syngas production, both CH₄ and CO₂ emissions into the atmosphere can be reduced. Because of its potential for reducing global warming, further understanding of syngas production from biogas is essential. Moreover, syngas can also be used for H₂ production. In this case, H₂ can be enriched via the water–gas shift reaction using syngas and H₂O as feedstock. Among the various alternative energy forms, hydrogen is considered an important energy carrier in the future [17]. It is also an important raw material in the chemical industry and can be used as a fuel in fuel cells to produce electrical energy. For reasons of sustainability, the use of renewable fuel sources such as biogas or biomass for hydrogen production has received considerable attention [18,19,20].

Several studies have reported on syngas production from biogas via the TR process experimentally. In the study of Vita et al. [21], tri-reforming simulated biogas over a Ni/ceria based catalysts was carried out and the H₂O/CH₄ and O₂/CH₄ molar ratios, reaction temperature, and nickel content effects on the catalyst’s performance were studied. They found that the H₂/CO ratio could be flexibly adjusted using added amounts of oxygen and steam in order to meet the requirements of downstream processes. In the study of Lau et al. [22], biogas was used as the fuel source in dry reforming and combined dry/oxidative reforming reactions. The gas stream temperature and reactor space velocity effects were examined experimentally. Their results indicated that an increase in the O₂/CH₄ ratio at low temperature promotes hydrogen production. In dry/oxidative reforming, they found that biogas dry reforming is dominant and the overall reaction is net endothermic when the reaction temperature is higher than 600 °C. In the study of Zhu et al. [23], biogas reforming with added O₂ through a spark-shade plasma was conducted under an O₂/ CH₄ ratio of 0.60 and CO₂/CH₄ ratios ranging from 0.17 to 1.00. Their results indicated that O₂ and CH₄ conversions decreased when the CO₂/CH₄ ratio was increased. They also reported that the partial oxidation of methane contributed mostly to CH₄ conversion and the reverse water–gas shift (WGS) reaction dominated in CO₂ conversion.

In addition to experimental work, several numerical TR process models using biogas as the feedstock have also been reported in the literature. In the study of Corigliano and Fragiacomo [24], biogas dry reforming analysis under various operating conditions was carried out using a numerical model. The CO₂/CH₄ ratio, pressure and temperature effects on reaction performance were reported. In the study of Hernández and Martín [25], a process based on mass and energy balances, chemical and phase equilibria, and rules of thumb was developed to optimize the production of methanol using biogas as the raw material. Based on the production cost and carbon footprint, the optimized CH₄/CO₂ ratio contained in the biogas was found. In the study of Hajjaji et al. [26], a H₂ production system via biogas reforming was investigated using life-cycle assessment (LCA). They found that the total GHG emissions from the system were about half of the life-cycle GHG of conventional H₂ production systems via steam methane reforming. In the study of Zhang et al. [27], the effects of various factors including reaction temperature, reactor pressure and CH₄ flow rate on the syngas compositions obtained from the TR process were investigated numerically. An optimum operating condition for syngas production with a target ratio and maximized CO₂ conversion were obtained.

In this work, the TR process is employed for syngas production using biogas as the feedstock. The effects of various operating conditions such as pressure, temperature, biogas composition, air addition, and H₂O additions are investigated. The novelty of this paper is the focus on CO₂ conversion in the TR process, which is seldom reported in the literature. Air is used as the added reactant in this study instead of pure oxygen in the conventional TR process.

2. Modeling

2.1. Chemical Reaction

The following reactions are coupled and carried out in a single reactor in the TR porcess:

Steam reforming (SR):

$${\mathrm{CH}}_4+{\mathrm{H}}_2\mathrm{O}\leftrightarrow \mathrm{CO}+3{\mathrm{H}}_2,\Delta {\mathrm{H}}_{298\mathrm{K}}^0=+206\mathrm{kJ}/\mathrm{mole}$$

(1)

Dry reforming (DR):

$${\mathrm{CH}}_4+{\mathrm{CO}}_2\leftrightarrow 2\mathrm{CO}+2{\mathrm{H}}_2,\Delta {\mathrm{H}}_{298\mathrm{K}}^0=+247\mathrm{kJ}/\mathrm{mole}$$

(2)

Partial oxidation (POX):

$${\mathrm{CH}}_4+0.5{\mathrm{O}}_2\leftrightarrow \mathrm{CO}+2{\mathrm{H}}_2,\Delta {\mathrm{H}}_{298\mathrm{K}}^0=-36 \mathrm{kJ}/\mathrm{mole}$$

(3)

As shown in Equations (1)–(3), the TR process combines the endothermic SR and DR reactions and the exothermic POX reaction. The heat released from POX is used as the heat supply for SR and DR and makes the TR process energy efficient [28]. As noted by Cho et al. [29], the chemical reactions involved in the TR process can be alternatively described using Equation (1) along with the following reactions:

Reverse CO₂ methanation (RCM):

$${\mathrm{CH}}_4+2{\mathrm{H}}_2\mathrm{O}\leftrightarrow {\mathrm{CO}}_2+4{\mathrm{H}}_2,\Delta {\mathrm{H}}_{298\mathrm{K}}^0=+165\mathrm{kJ}/\mathrm{mole}$$

(4)

Water-gas shift (WGS):

$$\mathrm{CO}+{\mathrm{H}}_2\mathrm{O}\leftrightarrow {\mathrm{CO}}_2+{\mathrm{H}}_2,\Delta {\mathrm{H}}_{298\mathrm{K}}^0=-41\mathrm{kJ}/\mathrm{mole}$$

(5)

Complete oxidation of methane (COM):

$${\mathrm{CH}}_4+2{\mathrm{O}}_2\leftrightarrow {\mathrm{CO}}_2+2{\mathrm{H}}_2\mathrm{O},\Delta {\mathrm{H}}_{298\mathrm{K}}^0=-803\mathrm{kJ}/\mathrm{mole}$$

(6)

Note that with the chemical reactions described in Equations (1) and (4)–(6), the TR process becomes the well-known catalytic partial oxidation of methane (CPOM). In the literature, there are many studies devoted to the analysis of kinetic mechanisms for CPOM [30,31]. Similar to other reforming process of CH₄, many reactions are likely to occur in the TR process. In addition to the study of Cho et al. [29], studies of De Groote and Froment [32], Scognamiglio et al. [33], Chan and Wang [34], and Izquierdo et al. [35] also reported that the reaction mechanism of CPOM is indirect in which the process can be described by combining reactions of methane oxidation, methane–steam reforming, and water–gas shift. According to these studies, reactions such as CO oxidation, H₂ oxidation, and the Boudouard reaction were not included.

Equations (1), (4) and (5) are the reactions involved in the conventional SR reaction. In this study, syngas under high pressure is of interest for further fuel synthesis. The kinetic model for the SR reaction over a nickel catalyst given by Xu and Froment [36] is adopted,

SR:

$${\mathrm{r}}_1=\frac{{\mathrm{k}}_1}{{\mathrm{p}}_{{\mathrm{H}}_2}^{2.5}}\left[{\mathrm{p}}_{{\mathrm{CH}}_4}{\mathrm{p}}_{{\mathrm{H}}_2\mathrm{O}}-\frac{{\mathrm{p}}_{{\mathrm{H}}_2}^3{\mathrm{p}}_{\mathrm{CO}}}{{\mathrm{K}}_{\mathrm{eq},1}}\right]/{\mathrm{DEN}}^2$$

(7)

WGS:

$${\mathrm{r}}_2=\frac{{\mathrm{k}}_2}{{\mathrm{p}}_{{\mathrm{H}}_2}}\left[{\mathrm{p}}_{\mathrm{CO}}{\mathrm{p}}_{{\mathrm{H}}_2\mathrm{O}}-\frac{{\mathrm{p}}_{{\mathrm{H}}_2}{\mathrm{p}}_{{\mathrm{CO}}_2}}{{\mathrm{K}}_{\mathrm{eq},2}}\right]/{\mathrm{DEN}}^2$$

(8)

RCM:

$${\mathrm{r}}_3=\frac{{\mathrm{k}}_3}{{\mathrm{p}}_{{\mathrm{H}}_2}^{3.5}}\left[{\mathrm{p}}_{{\mathrm{CH}}_4}{\mathrm{p}}_{{\mathrm{H}}_2\mathrm{O}}^2-\frac{{\mathrm{p}}_{{\mathrm{H}}_2}^4{\mathrm{p}}_{{\mathrm{CO}}_2}}{{\mathrm{K}}_{\mathrm{eq},3}}\right]/{\mathrm{DEN}}^2$$

(9)

$$\mathrm{DEN}=1+{\mathrm{K}}_{{\mathrm{CH}}_4}{\mathrm{p}}_{{\mathrm{CH}}_4}+{\mathrm{K}}_{\mathrm{CO}}{\mathrm{p}}_{\mathrm{CO}}+{\mathrm{K}}_{{\mathrm{H}}_2}{\mathrm{p}}_{{\mathrm{H}}_2}+\frac{{\mathrm{K}}_{{\mathrm{H}}_2\mathrm{O}}{\mathrm{p}}_{{\mathrm{H}}_2\mathrm{O}}}{{\mathrm{p}}_{{\mathrm{H}}_2}}$$

(10)

For COM, the kinetic model of Trimm and Lam [37] is adopted in this study,

COM:

$${\mathrm{r}}_4=\frac{{\mathrm{k}}_{4\mathrm{a}}{\mathrm{p}}_{{\mathrm{CH}}_4}{\mathrm{p}}_{{\mathrm{O}}_2}}{{(1+{\mathrm{K}}_{{\mathrm{CH}}_4}^{\mathrm{C}}{\mathrm{p}}_{{\mathrm{CH}}_4}+{\mathrm{K}}_{{\mathrm{O}}_2}^{\mathrm{C}}{\mathrm{p}}_{{\mathrm{O}}_2})}^2}+\frac{{\mathrm{k}}_{4\mathrm{b}}{\mathrm{p}}_{{\mathrm{CH}}_4}{\mathrm{p}}_{{\mathrm{O}}_2}}{(1+{\mathrm{K}}_{{\mathrm{CH}}_4}^{\mathrm{C}}{\mathrm{p}}_{{\mathrm{CH}}_4}+{\mathrm{K}}_{{\mathrm{O}}_2}^{\mathrm{C}}{\mathrm{p}}_{{\mathrm{O}}_2})}$$

(11)

Equation (11) was derived over Pt-based catalyst support, while the model adsorption parameters are adjusted for a Ni-based catalyst [38]. In Equations (7)–(11), ${\mathrm{r}}_{\mathrm{i}}$ is the reaction rate for SR (i = 1), WGS (i = 2), RCM (i = 3), and COM (i = 4); ${\mathrm{K}}_{\mathrm{eq},\mathrm{i}}$ and ${\mathrm{k}}_{\mathrm{i}}$ are the chemical equilibrium constant and rate constant for reaction i (i = 1,2,3,4); ${\mathrm{p}}_{\mathrm{j}}$ (j = CH₄, CO₂, H₂O, H₂, and CO) is the partial pressure of species j; and ${\mathrm{K}}_{\mathrm{j}}$ and ${\mathrm{K}}_{\mathrm{j}}^{\mathrm{C}}$ are the adsorption constants of species j. All of these kinetic parameters are given in the Arrhenius function type and are functions of temperature, and can be found in the literature [36,37]. It is noted that catalyst deactivation due to the thermal effect and carbon deposition is neglected in this study [39]. For a reforming reaction involving CH₄, carbon formation is inevitable. The carbon deposition on the catalyst surface is one of the reasons that causes catalyst deactivation. In the tri-reforming process, the appearances of O₂ and H₂O may suppress carbon formation [40,41]. Therefore, catalyst deactivation due to carbon deposition on the catalyst surface is neglected in this study.

2.2. Process Simulation

In this study, Aspen Plus (v.10) is employed to carry out the TR process using biogas as the feedstock. The flow process is depicted in Figure 1. The simulation is performed for a steady state. The biogas stream is assumed to be purely composed of CH₄ and CO₂ with the designated molar ratio. The air stream is composed of 21% O₂, 78% N₂, and 1% H₂. The purpose of H₂ addition is to avoid the singularity in chemical reaction rate computation. A 1% H₂ addition is determined through sensitivity analysis [42,43]. In order to produce high-pressure syngas for future use in fuel synthesis, two compressors (COM-1 and COM-2) are used to increase the biogas and air pressures. In the H₂O stream, a pump is used to increase the water pressure and it is then superheated in a boiler with heat supplied from the high-temperature product stream. After mixing in a mixer, the reactant mixture (TRI) is heated to a certain temperature before entering the insulated Rplug reactor (TR). The gas mixture from the reactor (TRO) is sent to the boiler where the heat is recovered for superheating the water. The TR process performance is characterized using the following dimensionless groups,

CH₄ and CO₂ conversions:

$${\mathrm{X}}_{\mathrm{i}}=\frac{{\mathrm{n}}_{\mathrm{i},\mathrm{in}}-{\mathrm{n}}_{\mathrm{i},\mathrm{out}}}{{\mathrm{n}}_{\mathrm{i},\mathrm{in}}}\times 100\%,\mathrm{i}={\mathrm{CH}}_4,{\mathrm{CO}}_2$$

(12)

H₂ yield:

$${\mathrm{Y}}_{{\mathrm{H}}_2}=\frac{{\mathrm{n}}_{{\mathrm{H}}_2,\mathrm{out}}-{\mathrm{n}}_{{\mathrm{H}}_2,\mathrm{in}}}{{\mathrm{n}}_{{\mathrm{CH}}_4,\mathrm{in}}}$$

(13)

CO yield:

$${\mathrm{Y}}_{\mathrm{CO}}=\frac{{\mathrm{n}}_{\mathrm{CO},\mathrm{out}}-{\mathrm{n}}_{\mathrm{CO},\mathrm{in}}}{{\mathrm{n}}_{{\mathrm{CH}}_4,\mathrm{in}}}$$

(14)

H₂/CO ratio:

$${\mathrm{H}}_2/\mathrm{CO}=\frac{{\mathrm{Y}}_{{\mathrm{H}}_2}}{{\mathrm{Y}}_{\mathrm{CO}}}$$

(15)

where ${\mathrm{n}}_{\mathrm{i},\mathrm{in}}$ is the molar flow rate of the i-th species supplied to the process; and ${\mathrm{n}}_{\mathrm{i},\mathrm{out}}$ is the molar flow rate of the i-th species at reactor outlet. Based on these definitions, CH₄ conversion is the ratio of the CH₄ consumption rate to the fed CH₄ flow rate at the reactor inlet. Similarly, CO₂ conversion is the ratio of the CO₂ consumption rate to the fed CO₂ flow rate at the reactor inlet. The H₂ and CO yields are defined as the net increased amounts of H₂ and CO from the reaction per fed CH₄ flow rate. The H₂/CO ratio is defined as the ratio of H₂ yield to CO yield. Note that all these variables are dimensionless.

In addition to the reactant conversion and product yield, energy and exergy analyses based on the resulting product stream flowing out of the TR reactor are also carried out. For the chemical reaction, there are several ways to define the first- and second-law efficiencies [44,45]. Since the objective of the TR process is to convert biogas into syngas, and noting that CO₂ has zero low heating value (LHV), the first-law efficiency is then defined as,

$${\mathsf{\eta }}_{\mathrm{I}}=\frac{{\mathrm{n}}_{\mathrm{CO},\mathrm{out}}{\mathrm{LHV}}_{\mathrm{CO}}+{\mathrm{n}}_{{\mathrm{H}}_2,\mathrm{out}}{\mathrm{LHV}}_{{\mathrm{H}}_2}}{{\mathrm{n}}_{{\mathrm{CH}}_4,\mathrm{in}}{\mathrm{LHV}}_{{\mathrm{CH}}_4}+{\mathrm{W}}_{\mathrm{comp}}+{\mathrm{W}}_{\mathrm{pump}}+{\mathrm{Q}}_{\mathrm{heat}}}\times 100\%$$

(16)

where ${\mathrm{W}}_{\mathrm{i}}$ (i = comp, pump) and ${\mathrm{Q}}_{\mathrm{heat}}$ are the input work and heat input, respectively. The main heat input occurs at the heater at which the mixed reactant is heated to a certain inlet temperature. The exergetic analysis is carried out by considering three exergy transfers:

Exergy due to work:

$${\mathrm{Ex}}_{\mathrm{W}}={\mathrm{W}}_{\mathrm{comp}}+{\mathrm{W}}_{\mathrm{pump}}$$

(17)

Exergy due to heat transfer:

$${\mathrm{Ex}}_{\mathrm{Q}}={\mathrm{Q}}_{\mathrm{heat}}(1-\frac{{\mathrm{T}}_0}{\mathrm{T}})$$

(18)

Exergy due to mass flow:

$${\mathrm{Ex}}_{\mathrm{f},\mathrm{i}}={\mathrm{N}}_{\mathrm{i}}\{\left[\left(\mathrm{h}-{\mathrm{h}}_0\right)-{\mathrm{T}}_0\left(\mathrm{s}-{\mathrm{s}}_0\right)\right]+\sum {\mathrm{x}}_{\mathrm{k}}{\mathrm{e}}_{\mathrm{k}}^{\mathrm{CH}}+{\mathrm{RT}}_0\sum {\mathrm{x}}_{\mathrm{k}}{\mathrm{lnx}}_{\mathrm{k}}\}$$

(19)

In these equations, the subscript 0 denotes the reference state (25 °C and 1 atm). The exergy due to mass flow is contributed by physical exergy, chemical exergy and mixing exergy as shown on the right-hand side of Equation (19). For second-law efficiency, this is generally defined as the ratio of exergy recovered to the exergy supplied,

$${\mathsf{\eta }}_{\mathrm{II}}=\frac{{\mathrm{Ex}}_{\mathrm{out}}}{{\mathrm{Ex}}_{\mathrm{in}}}\times 100\%$$

(20)

where ${\mathrm{Ex}}_{\mathrm{in}}$ and ${\mathrm{Ex}}_{\mathrm{out}}$ are the exergies supplied to and recovered from the system, respectively. Based on Figure 1, ${\mathrm{Ex}}_{\mathrm{in}}$ and ${\mathrm{Ex}}_{\mathrm{out}}$ are expressed as,

$${\mathrm{Ex}}_{\mathrm{in}}={\mathrm{Ex}}_{\mathrm{biogas}}+{\mathrm{Ex}}_{\mathrm{air}}+{\mathrm{Ex}}_{{\mathrm{H}}_2\mathrm{O}}+{\mathrm{Ex}}_{\mathrm{Q}}+{\mathrm{Ex}}_{\mathrm{W}},{\mathrm{Ex}}_{\mathrm{out}}={\mathrm{Ex}}_{\mathrm{product}}$$

(21)

3. Results and Discussion

The TR process using biogas as the feedstock is similar to the tri-reforming of methane (TRM). The only difference is the CH₄ and CO₂ composition. We developed this work from our previous study [46] and focused on using biogas as the feedstock. To verify the correctness of the model built in Aspen Plus, the TRM using the reactor geometry and reactant composition reported in the studies of Chein et al. [46] and Arab Aboosadi et al. [47] was carried out using the built model in Aspen Plus. Figure 2 shows the comparison between the temperature and gas species distributions predicted from a two-dimensional model [46] and from a model built in Aspen Plus. As shown in Figure 2, the agreements for both temperature distribution shown in Figure 2a and species mole fractions shown in Figure 2b are quite good at the reactor downstream. The discrepancies in the region near the reactor inlet zone is believed due to the difference between one- and two-dimensional modeling. Since the TR process performance is evaluated using results at the reactor outlet, good agreement between one- and two-dimensional results is expected. In addition to the comparisons between numerical models, experimental verification of the numerical model was given in our previous study [46]. Also note that O₂ is consumed rapidly as it enters the reactor shown in Figure 2b. That is, there will be no O₂ available for oxidation of CO or H₂ in the downstream of the reactor.

Based on the comparisons discussed above, the model built in Aspen Plus can be correctly extended to the TR process using biogas as the feedstock. The base operating conditions are listed in

Table 1. Reactor geometry and base operation conditions [46,47].

Parameter	Value
Reactor length, L	2 m
Reactor diameter, d(=2Rb)	10 mm
Inlet pressure, pin	20 atm
Inlet temperature, Tin	300~1000 °C
Reactant flow rate, F	0.0723 L min−1
Molar ratio of biogas CH4:CO2:Air:H2O	1:0.5:2:1
Catalyst	Ni/Al2O3
Catalyst size, dp	0.42 mm
Catalyst weight, W	0.25 g
W/F ratio	0.0576 ghL-1
Heat-transfer condition	Adiabatic

. The parameters listed in Table 1 are adopted from our previous study except for the feedstock composition and total volumetric flow rate [46]. The molar ratio of the reactants are chosen as CH₄:CO₂:Air:H₂O = 1:0.5:2:1 and the volume flow rate is fixed as 0.0723 L/min. As compared with the previous study [46], a higher volume flow is used in this study because of the presence of N₂ in the air. For economy, air is added instead of pure oxygen. The advantage of this is to avoid the cost of oxygen separation from air, but this obviously results in increased reactor volume. In the following, the TR process performance is examined using reactant inlet temperature T_in as the primary parameter. The effects of various pressures, catalyst weight/volume flow rate (W/F) ratios, CO₂/CH₄ ratio in biogas, amounts of air and H₂O on TR process performance are discussed.

Figure 3 shows the TR process performance using the base operations listed in Table 1. In Figure 3a the temperature variation along the reactor length for T_in = 900 °C is shown. Due to the methane oxidation reaction, the maximum temperature occurs in the near entrance region. The energy produced from methane oxidation is used for steam reforming and dry reforming in the reactor downstream. This causes the temperature to decrease along the reactor length. The CH₄ and CO₂ conversions are shown in Figure 3b. The abrupt increase in CH₄ conversion occurs at T_in = 550 °C. This indicates that T_in should be higher than 550 °C in order to activate the catalyst. With temperature higher than 550 °C, CH₄ conversion increases gradually with increased T_in. In Figure 3b, negative CO₂ conversion results for the low T_in regime. From the TR process chemical reactions, CO₂ is produced by the methane oxidation and WGS reactions and consumed by the dry reforming reaction. For low T_in, the WGS reaction is dominated and CO₂ consumption by DR is low. This results in negative CO₂ conversion. However, positive CO₂ conversion can result when T_in becomes higher than 700 °C, indicating that DR is active. DR contributes to increase the H₂ and CO yield in the high T_in regime. Figure 3b also indicates that a complex interaction between CO₂ consumption and production reactions occurs for T_in in the 500~600 °C range. The conversions of CH₄ and CO₂ from an equilibrium TR process obtained from an Aspen Plus simulation are also shown in Figure 3b using the parameters listed in Table 1. Since the results from the equilibrium process can be regarded as the theoretical limit of the reaction, it can be seen that CH₄ conversion from the catalytic reaction is lower than that from the equilibrium reaction. Due to more CO₂ production, lower CO₂ conversion results from the equilibrium reaction. For T_in higher than 800 °C, CO₂ conversion from the equilibrium reaction is higher than that from the catalytic reaction. From Figure 3c, the H₂ yield, CO yield and H₂/CO ratio are shown. It can be seen that when T_in is lower than 500 °C, the H₂ and CO yields are very low due to inactive catalytic reactions at low temperatures. In this low T_in regime, CO yield is much lower than H₂ yield and results in a high H₂/CO ratio. As T_in is higher than 550 °C, the H₂/CO ratio decreases with T_in slowly with a value close to 2. The decrease in H₂/CO with T_in is due to increased CO production from the DR reaction while H₂ decreases due to the reverse WGS reaction. In Figure 3d, the H₂ yield, CO yield, and H₂/CO ratio from the equilibrium TR process are also shown. As with conversions of CH₄ and CO₂ shown in Figure 3b, both yields of H₂ and CO from the catalytic reaction are lower than the equilibrium reaction. At high temperature, the H₂/CO ratio from both equilibrium and catalytic reactions is about the same. In Figure 3d, the first- and second-law efficiencies are shown. Based on Equation (16), the first-law efficiency depends on the H₂ and CO yields. Because of higher H₂ and CO yields at higher T_in, η_I increases with increased T_in. However, the variation in η_II is opposite that of η_I. Increased T_in implies that the chemical reaction is more complete towards the product side. Since the chemical reaction is a highly irreversible process, high exergy destruction due to the chemical reaction is expected. This results in decreased η_II as T_in increases. For the low T_in regime, exergy destruction due to the chemical reaction is low because of low catalytic activity. Moreover, the contributions of exergy destruction from compressors, pump, heaters and mixers are small. This leads to high η_II in the low T_in regime.

In Figure 4, the variation of species mole fraction of the TR process using the base operation conditions listed in Table 1 is shown. It can be seen that significant mole fraction variation can be found when T_in is higher than 500 °C. The mole fractions of reactants (CH₄, CO₂, H₂O, O₂, and N₂) decrease while the mole fractions of products (CO and H₂) increase as T_in increases. Due to a highly active methane oxidation reaction, O₂ is consumed completely when T_in is greater than 550 °C. Also note that the variation trend of mole fractions of CO and H₂ are similar to the yields of CO and H₂ presented in Figure 3c. The yields of H₂ and CO are used to characterize the TR process performance in this study.

In the following, parametric studies based on the base operation conditions listed in Table 1 are carried out. As listed in Table 1, the inlet temperature of the reactant is the primary parameter and the amount of CH₄ fed is used as the reference for the species contained in the reactant and product. The catalyst weight (W) is varied from 0.025 g to 2.5 g; the operation pressure (P) is varied from 10 to 30 atm; the air/CH₄ ratio is varied from 1 to 3; H₂O/CH₄ is varied from 1 to 3; and CO₂/CH₄ is varied from 0.25 to 0.75.

In Figure 5 the effect of W/F ratios on the TR process is examined. The results shown in Figure 4 were obtained by varying the catalyst weight, while other parameters listed in Table 1 were kept fixed. That is, higher W/F ratio results when the catalyst weight is increased. As shown in Figure 5a, a higher temperature along the reactor length is obtained for the W/F = 0.00576 ghL⁻¹ case. This indicates that a smaller amount of energy released from methane oxidation reaction is used for endothermic SR and DR reactions. For W/F = 0.0576 and 0.576 ghL⁻¹ cases, temperature variations are identical at the reactor downstream. That is, there is a limiting W/F ratio for the reaction. Increasing the W/F ratio (either increasing catalyst weight or decreasing reactant volumetric flow rate) may not lead to further improved reaction performance. In Figure 5b, CH₄ and CO₂ conversions are shown. Due to low catalyst activity, CH₄ conversion is low when T_in is low. It can be seen that the T_in at which CH₄ conversion abruptly increases can be decreased by increasing the W/F ratio. That is, the catalyst activation temperature can be lowered with increased W/F ratio. As shown in Figure 5b, the T_in at which CH₄ conversion increases abruptly are 700 °C, 500 °C and 400 °C for W/F = 0.00576, 0.0576, and 0.576 gLh⁻¹, respectively. CH₄ conversions for the W/F = 0.0576 and 0.576 gLh⁻¹ cases become identical when T_in is higher than 550 °C. As discussed above, limited performance results when the W/F ratio is increased. From Figure 5b, CO₂ conversion has a negative value except in the high T_in regime. This is due to CO₂ formation in the methane oxidation and WGS reactions while DR is less active. At high temperatures, CO₂ is consumed via the dry reforming reaction, leading to positive CO₂ conversion. In Figure 5c, the H₂/CO ratios for various W/F ratios are shown. It can be seen that H₂/CO ratio is about the same for the three W/F ratios studied when T_in is high. The H₂/CO ratio close to a value of 2 can be obtained for the W/F range studied. In Figure 5d, variations in η_I and η_II are shown. It can be seen that η_I can be enhanced by increasing the W/F ratio. However, η_II decreases when the W/F ratio is increased because of a more complete chemical reaction.

In Figure 6, the reactor operating pressure effect on TR process performance is examined. From Figure 6a the highest temperature increases with increased operating pressure. As the high temperature in the near-entrance region of the reactor is due to the methane oxidation reaction, this implies that methane oxidation can be enhanced by increasing the operating pressure. Due to the enhanced methane oxidation reaction, the T_in at which an abrupt increase in CH₄ conversion occurs can be decreased by increasing the pressure, as shown in Figure 6b. Figure 6b also shows that CH₄ conversion can be increased in the low T_in regime when the pressure is increased. That is, increased operating pressure can enhance catalyst activity at lower temperatures. In the high T_in regime, CH₄ conversion is slightly decreased as the pressure is increased. Although higher CH₄ conversion can be obtained from lower pressure operations, the resulting syngas may not be suitable for further use because most applications involve high-pressure synthetic processes. A H₂/CO ratio with a value close to 2 is obtained for all the pressures studied when T_in is high, as shown in Figure 6c. Because of the reduced CH₄ conversion at a high T_in regime, it can be seen that η_I decreases with increased T_in and pressure, as shown in Figure 6d. However, Figure 6d shows that η_II increases with decreasing pressure because of less exergy destruction by the chemical reaction.

The variation in reactant composition effect on TR process performance is examined in the following. Figure 7 shows the air amount effect. Figure 7a shows that temperature can be increased using more air as the reactant. That is, a more complete methane oxidation reaction is achieved when the air supply is increased. With the increase in air amount, both CH₄ and CO₂ conversions can be enhanced, as shown in Figure 7b. For the Air/CH₄ = 3 case, 100% CH₄ conversion can be reached for T_in higher than 550 °C. Due to the increased energy supply, dry reforming can occur in the lower T_in regime resulting in increased CO₂ conversion. However, negative CO₂ conversion is still found when T_in is low. Although more N₂ is also introduced, increasing the volumetric flow rate of the entire reactant, it does not affect CH₄ and CO₂ conversions. As shown in Figure 7c, a H₂/CO ratio with a value higher than 2 can be obtained for the Air/CH₄ = 3 case because DR is more active when the temperature is high. For the Air/CH₄ = 1 case, the H₂/CO value is lower than 2. This is due to the reverse WGS reaction at high temperatures, reducing the H₂ amount. Because of decreased H₂ yield, lower η_I in the higher T_in regime is obtained, as shown in Figure 7d. The reverse WGS reaction also causes η_II to increase with T_in in the high T_in regime.

Figure 8 shows the H₂O amount effect on TR process performance. With increased H₂O in the reaction, lower temperature results at the reactor entrance region, as shown in Figure 8a, because of an increased reactant volumetric flow rate and endothermic SR reaction. The increased H₂O amount does not affect CH₄ conversion, as shown in Figure 8b. However, more negative CO₂ conversion results. In addition to CO₂ produced from methane oxidation, CO₂ may also be produced from WGS and RCM reactions, as indicated in Equations (4) and (5) when H₂O is increased. As shown in Figure 8c, a higher H₂/CO ratio is obtained when H₂O is increased because of increased H₂ yield. Figure 8d shows lower η_I results when the H₂O amount is increased. This is because higher heating to the reactant is required when the H₂O amount is increased. η_II increases with increased H₂O amount, indicating that less exergy destruction results as H₂O is increased.

Figure 9 shows the amount of CO₂ contained in the biogas effect on the TR process. As shown in Figure 9a, the amount of CO₂ does not affect the reaction temperature to a large extent. The temperature increases slightly as the CO₂ amount is decreased. As shown in Figure 9b, CH₄ conversion is affected insignificantly by the CO₂ amount. However, CO₂ conversion is always negative for the CO₂/CH₄ = 0.25 case. That is, more CO₂ is produced as a result of SR and WGS reactions than that consumed by DR and reverse WGS reactions. In Figure 9c, higher H₂/CO results when CO₂ is decreased. This may be due to less CO formed from CO₂ conversion. As with CH₄ conversion, the CO₂ amount effect on first- and second-law efficiencies is not significant, as shown in Figure 9d.

Typical H₂ and CO yield results are shown in Figure 10 for various air and H₂O amounts. In Figure 10a, the H₂ and CO yields increase with the increased air added in the reactant. As a high temperature results in the Air/CH₄ = 3 case, the H₂ yield decreases with increased T_in due to the reverse WGS reaction. Figure 10b shows that H₂ yield can be enhanced by increased H₂O addition. CO yield also decreases with H₂O addition because of inactive DR and reverse WGS reactions. As a result, a higher H₂/CO ratio is obtained, as shown in Figure 8b.

From the results shown above, the CO₂ conversion is low or negative (CO₂ production). Positive CO₂ conversion only occurs in the high T_in regime. It is, then, desirable to determine the way to enhance CO₂ conversion in the TR process. After several sets of numerical experiments, it was found that high CO₂ conversion can be obtained when H₂O is low. In this case, the TR process approaches the dry reforming of methane (DRM). Figure 11 shows TR process performance with H₂O/CH₄ = 0.001. As shown in Figure 11a, the temperature drop occurs in the region very near the entrance because DRM is a highly endothermic reaction. Large amounts of required heat leads to this temperature drop. When the methane oxidation becomes active, energy release causes a temperature increase in the reactor downstream. In Figure 11b, CH₄ and CO₂ conversions are shown for the H₂O/CH₄ = 0.001 case. CO₂ conversion is always positive and increases with increased T_in. The CO₂ conversion is lower than that of CH₄ because of a low CO₂/CH₄ ratio in the biogas. Because of small amounts of H₂O, the H₂/CO ratio is close to unity, which is the stoichiometric H₂/CO ratio of DRM, as shown in Figure 11c. Figure 11d shows that η_I increases with increased T_in because of higher H₂ and CO yield. The second-law efficiency states that η_II decreases with increased T_in because of higher exergy destruction when the chemical reaction is more complete.

4. Conclusions

The tri-reforming process was used in this study for syngas production from biogas. The effects of various operating parameters such as pressure, temperature and reactant composition were studied based on a model built in Aspen Plus. Based on the results obtained, the following conclusions can be drawn:

(1)

There appears to be a limiting space velocity for the reaction. Beyond this limiting value, the reaction approaches the same performance. Lowering the reaction pressure could lead to higher CH₄ conversion, but the syngas produced may not be suitable for further applications.

(2)

CH₄ and CO₂ conversions can be enhanced by increasing the amount of air in the reactant. Higher amounts of air could result in decreased H₂ yield due to the reverse water–gas shift reaction, which is favorable at high reaction temperatures.

(3)

A higher H₂/CO ratio can be obtained by increasing H₂O addition. However, the dry reforming reaction is suppressed, leading to low CO₂ or negative conversion.

(4)

Dry reforming of CO₂ can only be found when the reaction temperature is high. This results in positive CO₂ conversion and contributes to increased H₂ and CO yields.

(5)

Higher CO₂ conversion can be obtained for the low H₂O addition case. However, low H₂/CO with a value close to unity results.

(6)

The first-law efficiency increases with the increased reaction temperature because of higher H₂ and CO yields. The second-law efficiency decreases with the increased temperature because of higher exergy destruction due to a more complete chemical reaction at high temperatures.