A polygeneration process concept for HCCI-engines – Modeling product gas purification and exergy losses

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A polygeneration process concept for HCCI-engines e Modeling product gas purification and exergy losses R. Hegner*, B. Atakan Institute for Combustion and Gas Dynamics e Thermodynamics, University of Duisburg-Essen, 47048 Duisburg, Germany

article info

abstract

Article history:

This modeling study addresses the question, whether HCCI piston engines can be used as

Received 8 July 2016

chemical reactors in a polygeneration process. In this context, a combustion engine is used

Received in revised form

to produce hydrogen by partial oxidation in the engine and the further auxiliary units

5 September 2016

assure the purification of hydrogen. Several aspects were addressed during the develop-

Accepted 8 September 2016

ment of the process regarding fresh-gas- and exhaust gas-treatment and energy integra-

Available online xxx

tion. The fresh gas preheating for methane ignition was achieved by recirculation of hot exhaust gas, which also contributed to a high flexibility towards power-, heat- and

Keywords:

hydrogen-output. Exergetic efficiencies, process-outputs and fuel consumption were

Polygeneration

calculated and compared for different operation points. Exergetic efficiencies of up to 80%

HCCI

were achievable and this high efficiency leads to reduced fuel consumption, with up to 40%

Exergy

fuel savings compared to the separated production of power, heat and hydrogen. Power

EGR

and heat flow of the process can be adjusted very flexibly and the ratio can be varied within a factor of two within the investigated operating conditions. © 2016 Hydrogen Energy Publications LLC. Published by Elsevier Ltd. All rights reserved.

Introduction These days energy and process engineering rely heavily on the conversion of fossil energy sources. Due to the limited availability of these resources, improved efficiency and flexibility of energy conversion is needed. With respect to the integration of strongly fluctuating renewable energies, flexibility in energy conversion gains importance in order to adjust the electrical energy generation to the actual demand. However most engines and turbines suffer from low efficiencies, especially under partial load conditions [1,2]. An option to achieve this flexibility is known as polygeneration, which is the simultaneous generation of work, heat and chemicals.

Usually polygeneration is discussed as the combination of processes, without changing the processes itself [3]. Wang et al. [4] investigated a polygeneration process that combines acetylene production and fuel cells. Here acetylene is produced by a partial oxidation reactor and the by-product synthesis gas is used in a fuel cell, while acetylene is separated via absorption. The exergetic efficiency of this combined process was found to be 43% and therefore more than 20% higher than the conventional acetylene production process without fuel cell. Polygeneration systems usually intend to provide a welladjusted combination of various outputs, thus lacking flexibility. Often coal or biomass is the primary energy source [5], but also complex fuel sources including waste streams are considered [6,7]. High temperature fuel cells are also one

* Corresponding author. Fax: þ49 203 379 1594. E-mail address: [email protected] (R. Hegner). http://dx.doi.org/10.1016/j.ijhydene.2016.09.050 0360-3199/© 2016 Hydrogen Energy Publications LLC. Published by Elsevier Ltd. All rights reserved. Please cite this article in press as: Hegner R, Atakan B, A polygeneration process concept for HCCI-engines e Modeling product gas purification and exergy losses, International Journal of Hydrogen Energy (2016), http://dx.doi.org/10.1016/j.ijhydene.2016.09.050

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possibility to convert methane or natural gas to electrical energy and hydrogen [8]. Cormos [9] developed a combined polygeneration process for power, hydrogen and carbon dioxide production, consisting of a bio-methanol steam reformer, carbon capture technologies and a steam turbine. He investigated different designs regarding how and where carbon dioxide is separated. In this case chemical looping concepts proved to be advantageous over amine-based absorption technologies. A combination of a fuel cell with absorption cooling was investigated by Yu et al. [10]. The process intended to provide electricity together with district heating or cooling. An exergetic efficiency of up to 84% was achieved in this study. In contrast to this conventional polygeneration approach, where separated processes are combined, increased flexibility could be achieved by using an internal combustion (IC) engine at fuel rich operation conditions as a kind of chemical reactor. Such a polygeneration engine would be able to flexibly change between various ratios of mechanical work, process heat, and chemicals, ranging from syngas (a mixture of H2 and CO) to unsaturated or partially oxidized hydrocarbons. The present work is focused on the production and purification of hydrogen by partial oxidation in a homogeneous charge compression ignition (HCCI) engine. Due to the homogeneous charging of the engine with a premixed fuel oxidizer mixture and the homogeneous self-ignition due to compression, the operation is mainly controlled by chemical kinetics, while flame propagation and fluid dynamics play a minor role, facilitating the concept evaluation. Operating such an engine with a lack of oxygen leads to partial oxidation, decreasing the work output compared to stoichiometric conditions. At moderately rich conditions the main partial oxidation products would be syngas. In contrast to lean combustion, a part of the fuel exergy remains in the product gas, which can be used either for other purposes in chemical industry or can be stored and burned later. This makes this kind of polygeneration process especially suitable for adapting the electrical power production to the demand. Syngas generation in engines has been studied by Karim et al. [11] with mixtures of CH4/O2-enriched air and pilot ignition by adding diesel fuel. Yang et al. investigated HCCI combustion in rich CH4/O2 mixtures with added CO and were successful in syngas production, while Szezich investigated a spark ignition (SI) process with CH4 at low compression ratios [12,13]. Morsy modeled an HCCI process with methane to find some good conditions for syngas production [14] and proposed to start compression at relatively high temperatures to achieve ignition. McMillan and Lawson experimentally investigated a fuel-rich natural-gas SI process, but also modeled an HCCI process [15]. With SI, they were successful in producing syngas up to a fuel-to-air equivalence ratio 4 ¼ 1.62. Just recently Lim et al. [16] investigated a SI-engine experimentally, operating at equivalence ratios up to 2.8. The authors discovered, that the H2-to-CO-ratio and the fuel conversion are increased by hydrogen and ethane addition to the fuel. Elevated equivalence ratios and delayed spark ignition timing also increased the H2to-CO-ratio but lead to reduced methane conversion. Exhaust soot concentrations raised at equivalence ratios above 2.4, but were considerably lower below this threshold. Gossler and Deutschmann [17] applied an optimization model similar to the one used in this work to investigate the hydrogen yield of a

fuel rich operated HCCI engine. Optimal values for pressure, engine speed and equivalence ratio were found dependent on the inlet temperatures. Thus, the hydrogen yield could be increased from 75% to 90% throughout the considered temperature range, with a strong linear relation towards the initial temperature. However, most of these publications concentrate on syngas yields and operation parameters but are not analyzing the overall thermodynamics. Investigating the mere production of synthesis gas or hydrogen in a combustion engine is also limited, since the product gas is a mixture of syngas with water, carbon dioxide and nitrogen byproducts, which need to be removed. Here, we are concentrating on hydrogen as the chemical product of the polygeneration process, and it is obvious that a purification with additional processes will be needed in order to use the hydrogen technically. Only when such a total polygeneration process is designed a comparison with conventional alternative systems like syngas production systems with steam reforming and water gas shift is meaningful. Thus, an overall process concept is designed here, that includes the central polygenerating engine, as well as necessary auxiliary units, like preheaters or separation units. In order to identify optimal operation points, a parameter study is carried out. Exergy [18], also known as available energy, is known to be a non-conserved quantity which may be used to assess polygeneration processes, which are usually designed to have low exergy losses. The higher the exergy loss of a process is, the more it deviates from a thermodynamically ideal, meaning reversible, process. The overall exergy loss including chemical and physical exergies of such processes is calculated here. The crank angle dependent, and thus time dependent, chemical conversion within the engine is calculated using chemical elementary reaction mechanisms within a single-zone HCCI model. Together with the second law of thermodynamics the exergy loss within the engine is calculated. In addition, the exergy losses in all other process units are evaluated. Methane was investigated as the fuel and dry air as oxidizer. The usage of methane makes the present work comparable to several previous studies [19], but using a fuel consisting of 95% methane, 4% ethane and 1% propane was also tested for some conditions. This led to slight changes in quantitative results but did not influence the overall picture. It turns out, that the needed initial temperature with typical engine speeds and compression ratios are quite high in order to obtain chemical conversion. Therefore fresh gas preheating in combination with exhaust gas recirculation is evaluated to elevate the inlet temperatures and to reduce peak pressures during ignition.

Modeling The in-cylinder processes were simulated with timedependent chemical kinetics considering the compression and the expansion stroke, starting at bottom dead center with closed valves. Modeling of selected auxiliary units will be described in Section “process concept” in more detail. The calculations were performed using the framework of the reactor model in Cantera within Python [20]. A homogeneous gas mixture was assumed for each state. Reactions were

Please cite this article in press as: Hegner R, Atakan B, A polygeneration process concept for HCCI-engines e Modeling product gas purification and exergy losses, International Journal of Hydrogen Energy (2016), http://dx.doi.org/10.1016/j.ijhydene.2016.09.050

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simulated with a detailed reaction mechanism from USC (111 species, 784 reactions) [21]. The engine model developed here, including wall heat transfer, is comparable to the single-zone model by Caton and Zheng [22]. Engine parameters are listed in Table 1 and were chosen according to typical automotive engine parameters. The somewhat high engine speed of 3000 rpm is associated with the intended high hydrogen production and engine power, as well as low heat losses. The model divides the closed engine cycle into small time steps. For each time step the following set of differential equations describing the conservation of species and energy are solved simultaneously by Cantera. dni ¼ Vu_ i ; dt

(1)

with the effective rate of formation u_ i of a species i, calculated by the sum of all rates of reactions j that contribute to the formation or consumption of this species. u_ i ¼

NR X

NS  00 Y y0 cijij kj yij  y0ij

j

! ;

(2)

i

where the total number of reactions and species are NR and NS respectively. The stoichiometric coefficients of species i in 00 reaction j are referred to as n for products and n0 for reactants. The conservation of energy is described in Eq. (3), where Hi stands for the molar enthalpy of species i, R is the universal gas constant, T the temperature and p the pressure. n_i represents the temporal change of moles, calculated with Eq. (1) and cv the mean, molar heat capacity at constant volume. dT ¼ dt

P ðn_i Hi  n_i RTÞ  pV_  Q_ 1

(3)

cv

The temporal change of the cylinder volume V_ at each time step was calculated from the piston kinematics determined by the engine geometry [23]. The time-dependent piston velocity is calculated from Eq. (4): ! p cosðqðtÞÞ ~* *sinðqðtÞÞ* 1 þ ; uðqðtÞÞ ¼ u 2 1=2 2 ðC2  sin ðqðtÞÞÞ

(4)

~, the constant C ¼ 3.5 [23] and the with the mean piston speed u time dependent crank angle, determined by Eq. (5): qðtÞ ¼ 2pNt þ p

(5)

The heat flow Q_ was estimated using Woschni's semiempirical formulas for the convective heat transfer coefficient [24]. At the end of the expansion stroke, the product-gas composition, entropy generation Sirr , and total work W where calculated, the latter from evaluating the integral

Table 1 e Engine properties and operation parameters. Parameter Displacement VD Bore d/stroke s Comp. ratio ε Engine speed N Intake pressure p0 Coolant temperature TK

Specification

3

V Z360

W¼

pdV

(6)

V0

along the cycle. Beyond the work, since an incomplete oxidation is performed, also a considerable amount of (chemical) exergy is present in the product gases. The product gas in rich combustion for polygeneration contains chemically useful species and unconverted chemicals and a considerable thermal internal energy. The exergy losses El are calculated using the Gouy-Stodola theorem (El ¼ Tsur Sirr ) [18] with the ambient temperature Tsur ¼ 298 K and the entropy generation Sirr of the process. The corresponding exergetic efficiency hex is determined with Eq. (7) using the fuel mass mFuel assuming that the exergetically valuable outputs (work, heat, and hydrogen) are used outside the process and not lost to the environment. The specific chemical exergy of methane efuel ¼ 52.281 MJ/kg was taken from Ref. [25], while the exergy of air is negligible and was set to zero. hex ¼ 1 

El mfuel efuel

(7)

The irreversible entropy production of each process unit Sirr is calculated by the second law or thermodynamics for open, steady systems [18]: Sirr ¼

X

mout *sout 

X

min *sin 

Q Tref

(8)

, where min and mout are the respective masses entering and exiting the unit and Q is the heat transferred. Two cases were analyzed regarding the reference temperatures Tref which is the assumed temperature to which heat is transferred from the process: The first case is that the transferred heat of the respective unit Q is used in other processes, according to the polygeneration concept. In this case, the reference temperature is set to 100 K below the mean temperature of heat exchanger, with Tmean defined as: Tmean ¼

hout  hin sout  sin

(9)

The second case is that the heat is not used and therefore it is lost to the surroundings. Here the Tref is set to the ambient temperature of 298 K. With the calculations and modeling described in this section, the product gas composition of the engine, and exergy losses of individual units are calculated. The next section describes the challenges and considerations of merging the engine and several auxiliary units into a functioning and flexible polygeneration process concept.

Process concept Preheating temperatures and product gas composition

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400 cm 79.5/80.5 mm 16.5:1 3000 RPM 101.3 kPa 373 K

The combustion engine is the central unit of the polygeneration approach presented here. Investigating the engine itself, grants some insight into the requirements crucial for a process concept. Since methane is a quite inert fuel, it is expected, that the fuel conversion is low at low temperatures. HCCI

Please cite this article in press as: Hegner R, Atakan B, A polygeneration process concept for HCCI-engines e Modeling product gas purification and exergy losses, International Journal of Hydrogen Energy (2016), http://dx.doi.org/10.1016/j.ijhydene.2016.09.050

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engines rely on auto-ignition of a homogeneous, premixed charge, so temperatures at the top dead center need to be high enough to overcome the activation energies and accomplish a sufficiently high radical concentration. This reasoning leads to relatively high necessary intake temperatures, in order to achieve ignition within 10 crank angle (CA) after top dead center (TDC). The latter constraint was mainly chosen, to avoid ignition during the compression strokes, which leads to high pressure rise rates which may lead to engine damage and also leads to heat losses especially under HCCI conditions. Although ignition later than 10 after TDC is also possible, the fuel conversion under these condition was found to be considerably lower (below 80%) compared to ignition between 0 and 10 after top dead center. Fig. 1 shows the crank-angle resolved temperature- and CH4-profiles for intake temperatures T0 of 620 K, 660 K and 700 K at an equivalence ratio of 4 ¼ 2:5 to illustrate this issue. 260 data points (one for each crank angle) are depicted in this figure but for clarity markers are only set every 20 crank angle. At T0 ¼ 620 K no ignition takes place and methane conversion is small. At 660 K intake temperature the ignition happens in the right time frame of around 10 after TDC and at even higher temperatures of 700 K, ignition already starts in the compression stroke, which is also not favorable due to higher heat losses and possible engine damage. Similar investigations were performed for various other equivalence ratios in order to identify required intake temperatures for proper timed ignition. Results for different equivalence ratios are shown in Fig. 2. The required intake temperatures for ignition and conversion are increasing with the equivalence ratio. This is due to the high heat capacity of methane compared to the one of air. Therefore, an increasing fuel content in the fresh gas mixture lowers the temperature after compression. Thus, intake temperatures need to be adjusted to the equivalence ratio. Kinetic effects due to the decreasing O2-concentration

Fig. 1 e Kinetically calculated temperature- and CH4-profile within HCCI cycles at 4 ¼ 2.5 for intake temperatures of 620e700 K, all other parameters like compression ratio, engine speed etc. were given in the text.

with increasing 4 may also contribute to the increasing necessary intake temperature, but the temperature at which ignition starts was always near 1300 K and thus comparable for each investigated equivalence ratio. Therefore, kinetic effects seem to play a minor role only. For polygeneration fuel rich equivalence ratios of at least 4 ¼ 2 are needed leading to initial temperatures above 600 K. From these calculations also all thermodynamic states along the cycle were calculated, including the gas composition at the end of the expansion stroke. For the fresh gas conditions from Fig. 2 the mole fractions of the major product gas species are shown in Fig. 3 as a function of equivalence ratio. In this diagram it is seen that the mole fractions of hydrogen and carbon monoxide have their respective maxima at an equivalence ratio of 4 ¼ 2:5. Lower equivalence ratios lead to undesired further oxidation of these product species to water and carbon dioxide. At higher equivalence ratios the fuel conversion and reaction rates are limited by the low combustion temperatures of about 1600e1900 K, but also obey the deficiency of oxygen, leading to lower mole fractions of hydrogen and carbon monoxide. However, high equivalence ratios also favor the generation of hydrocarbons like acetylene (C2H2) with mole fractions of up to 2.5%, which may prove to be a desirable product species of the polygeneration process in further investigations. Unfortunately, hydrocarbon formation during combustion is often coupled to soot formation as well. However, by focusing on hydrogen as target chemical and using equivalence ratios below 2.5, the maximal combustion temperatures usually reach values around 2200 K. At these temperatures and equivalence ratios soot is still mostly oxidized during the expansion stroke as Merker et al. [26] and Lim et al. [16] concluded. Admittedly, soot formation itself was not investigated in this work. Besides hydrogen, the product gas of the polygeneration engine also contains several species which can be regarded as unwanted byproducts, mainly water and carbon dioxide but also high amounts of nitrogen, which is not shown in Fig. 3. Table 2 lists the mole fractions of important product gas species at an equivalence ratio of 4 ¼ 2.5.

Fig. 2 e Required intake temperatures as a function of equivalence ratio for ignition in a HCCI process within 10 CA after TDC.

Please cite this article in press as: Hegner R, Atakan B, A polygeneration process concept for HCCI-engines e Modeling product gas purification and exergy losses, International Journal of Hydrogen Energy (2016), http://dx.doi.org/10.1016/j.ijhydene.2016.09.050

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Fig. 3 e Product gas composition at different fresh gas equivalence ratios at the end of an HCCI cycle. All other parameters except the equivalence ratio and initial temperature (see Fig. 2) were kept constant and are given in the text.

These considerations result in two major requirements which a process concept should meet. On the fresh gas side of the engine, the temperature of the fresh gas needs to be increased from ambient temperature to around 600 K (see Fig. 2) to assure ignition in the HCCI-engine. This preheating could be achieved by either heat exchanger using the thermal energy of the exhaust gas or by partial recirculation of the hot exhaust gases. The next section will reveal, that a combination of both, preheater and exhaust gas recirculation (EGR) is the most promising alternative. On the product gas side of the engine, the target species hydrogen should be separated from the byproducts and the remaining exhaust gas prepared for recirculation. Additional units or reactors, which increase the hydrogen yields and using the exergy of the other byproducts would also be preferable. Based on these considerations, a process concept was designed which includes the fresh gas preheating and the purification of hydrogen.

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cycles. After one cycle is completed the respective exhaust gas is partially fed back to the engine and the next cycle is calculated with new inlet gas composition. This procedure is repeated until a constant product gas composition is achieved to determine the steady state solution. This is usually accomplished within 3e4 cycles. Besides preheating, another purpose of this exhaust gas recirculation is the flexible adaption of engine power and hydrogen output to the demand, since increased exhaust gas contents result in less fuel introduced into the engine. Therefore, to gain high flexibility a wide range of utilizable exhaust gas contents is preferable. In order to quantify the amount of recirculated exhaust gas, the unitless exhaust gas ratio xEGR is introduced. It is defined as the molar ratio of the amount of recirculated exhaust gas nEG, to the total amount of substance (mol) introduced into the engine, consisting of the amount of exhaust gas nEG and the amount of fresh gas nFG: xEGR ¼ nEG =ðnFG þ nEG Þ

(10)

This relative amount of exhaust gas is varied in a parametric study, to investigate its effect on the ignition characteristics. The results for fuel conversion and crank-angle resolved temperature profiles for different amounts of recirculated exhaust gas are shown in Fig. 4 and Fig. 5. The temperature of the fresh gas is assumed to be 298 K, only for these calculations which were carried out to evaluate the exhaust gas ratio xEGR at which methane conversion would start without any additional preheating. The temperature of the exhaust gas for the results shown in Fig. 4 was fixed to 723 K, which is explained later in this section in more detail. In Fig. 4 a rapid increase in fuel conversion is noticed above 55% exhaust gas ratio. At lower xEGR the fuel conversion is nearly zero and at xEGR > 55% the fuel conversion reaches approximately 100%. The cause of this trend can be derived from the crank-angle resolved temperature- and CH4 conversion profiles in Fig. 5. At comparably low xEGR ¼ 40% the exhaust gas content and resulting temperature at the top dead

Process concept The preheating is partially accomplished by recirculating a part of the exhaust gas mixture at high temperature back into the engine. However, it should be mentioned, that the product hydrogen is separated from the exhaust gas before recirculation to maximize the hydrogen yield. So the composition and temperature of the gas fed back to the engine is not the same as the actual exhaust gas exiting the engine. For easier readability, in the following sections we will, however, still refer to this gas as exhaust gas. The exhaust gas recirculation is included in the model by calculating consecutive engine

Table 2 e Product gas composition at 4 ¼ 2.5 Species N2 H2 CO H 2O CO2

Mole fraction/% 51.1 23.2 14.9 9.3 1.5

Fig. 4 e Fuel conversion in an HCCI process with different amounts of exhaust gas xEGR ¼ nEG/(nFG þ nEG) with an intake equivalence ratio of 4 ¼ 2.5. These calculations were carried out for constant initial temperatures and exhaust gas temperatures (see text).

Please cite this article in press as: Hegner R, Atakan B, A polygeneration process concept for HCCI-engines e Modeling product gas purification and exergy losses, International Journal of Hydrogen Energy (2016), http://dx.doi.org/10.1016/j.ijhydene.2016.09.050

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εMembr ¼

Fig. 5 e Crank-angle resolved temperature profiles (solid lines) and methane conversion (dashed lines, right scale) at different xEGR ¼ nEG/(nFG þ nEG) at with an intake equivalence ratio of 4 ¼ 2.5. These calculations were carried out for constant initial temperatures and exhaust gas temperatures (see text).

center are still too low to enable ignition, which leads to low fuel conversion as well. At the ideal exhaust gas content xEGR ¼ 55% the temperature after compression is just right to assure ignition during the expansion stroke, at 185 CA, leading to around 10% unconverted fuel. Higher xEGR's shift the ignition into the compression stroke, which results in nearly complete fuel conversion. However, the high pressure rise rates during ignition, combined with the pressure rise due to compression can cause engine damage. This also increases the work needed for compression and may lead to higher heat losses. Therefore, merely mixing exhaust-gas and fresh-gas without any pretreatment would impose high restrictions on the process with regards to flexibility, since only xEGR > 55% would be technically feasible. This is avoided by the combined use of exhaust gas recirculation and heat exchangers, where thermal energy is transferred from the exhaust gas to preheat the fresh gas to elevate the temperatures to values above 298 K without changing the composition. At low xEGR's below 55% an additional preheater is used to elevate the fresh gas temperatures to the required level before mixing it with the exhaust gas. At higher xEGR's above 55% the preheater is not necessary anymore, but instead an exhaust gas cooler reduces the exhaust gas temperature, to shift the ignition time towards the expansion stroke. The purification of the product hydrogen may be achieved by either separating the byproducts or the hydrogen itself from the exhaust gas mixture. Since the hydrogen mole fraction is usually around 23% (see Table 2), it is reasonable to separate the hydrogen instead of the higher amounts of byproducts. Membranes for hydrogen separation are now in a high development stage with high selectivity and effectiveness at high temperatures and pressures [27]. The membrane effectiveness εMembr is defined as the molar amount of inlet hydrogen nH2 ;in that is found in the product stream nH2 ;Prod , see Eq. (11).

nH2 ;Prod nH2 ;in

(11)

Adhikari and Fernando [28] give a detailed overview of hydrogen membrane separation techniques. Usually metallic or ceramic membranes are operated at trans-membrane pressures of 1e40 bars and temperatures of 573e873 K and 873e1173 K respectively. Here a metallic membrane with εMembr ¼ 0:9 was assumed, according to typical values from Ref. [28]. This makes them well suited for separating hydrogen from the hot engine product gases. The resulting process concept with all auxiliary units is shown in Fig. 6. At the beginning methane and air are mixed at ambient temperature (state 1). The fresh-gas equivalence ratio 4FG defines how much methane is mixed with air. The fresh gas is then led into an optional preheater and heated to the fresh gas temperature T2 (state 2). The preheater is only necessary if the fresh gas is afterwards mixed with less than 55% recirculated exhaust gas, with the respective temperature T12 (state 12). This mixture is than led into the polygeneration engine with the temperature T3 (state 3). The combustion inside the HCCIengine is simulated with the engine model and the product gas leaves the engine at the calculated final temperature T4 (state 4). Afterwards it flows through two heat exchangers. The heat transferred in cooler 1 can be used as process heat of the polygeneration process, if a suitable coolant is available. However specific properties of secondary fluids in all coolers are neglected in order to reduce the degrees of freedom. Instead of that, the process heat is calculated according to the first law of thermodynamics and the exergy loss due to heat transfer is either accounted with a mean temperature difference of 100 K, as described in Section “modeling”, or the transferred heat (if not used) is regarded as exergy loss. If the heat is used outside the process a small temperature difference between exhaust gas and coolant is thermodynamically preferable. Here, the mean temperature difference is set to 100 K to allow small heat exchange areas. Both cases will be discussed in Section “exergetic efficiency”. After cooler 1 (state 5) further heat is provided to the steam generator, which evaporates liquid water at the exhaust gas pressure and 298 K to supply steam to the downstream water-gas-shift reactor. This reactor is used to increase the hydrogen yield, by converting carbon monoxide, exhausted by the engine, catalytically with steam into hydrogen, according to Eq. (12). CO þ H2 O #H2 þ CO2

(12)

The molar steam-to-carbon-ratio is the molar steam flow rate (n_s ) entering the water-gas-shift reactor divided by the molar carbon monoxide flow rate (n_co ), contained in the exhaust gas. The value was chosen to be nn__cos ¼ 1:3, to achieve sufficient carbon monoxide conversion at the chosen equilibrium temperature and was calculated as average from two publications on steam-methane-reforming (SMR) [19,29]. This reaction, as well as the subsequent hydrogen separation is carried out at 723 K and the respective exhaust gas pressures. The chosen temperature is a compromise to meet two restrictions: metallic hydrogen membranes operate best at high temperatures of around 673e873 K [28] and the exothermal W/ G-shift-reaction, which has highest conversion rates at low temperatures 473e573 K, but also a reasonable conversion

Please cite this article in press as: Hegner R, Atakan B, A polygeneration process concept for HCCI-engines e Modeling product gas purification and exergy losses, International Journal of Hydrogen Energy (2016), http://dx.doi.org/10.1016/j.ijhydene.2016.09.050

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Fig. 6 e Polygeneration process concept for HCCI-engines with H2 as chemical product. The preheater is used at small xEGR < 55%, the cooler 3 is used at larger xEGR > 55%.

rate at 723 K [29]. Chemical equilibrium under isobaric and isothermal conditions is assumed for the water-gas-shift reaction (7) and was calculated using the appropriate thermodynamic data. Mole fractions of the involved species at chemical equilibrium (state 7) were determined based on their initial mole fractions and the extent of reaction. The hydrogen membrane is modeled as isothermal, metallic membrane with a typical prescribed molar effectiveness of εMembr ¼ 0:9 providing 100% pure hydrogen (state 8) [28]. The separated hydrogen is then cooled to the ambient temperature (state 9) by the cooler 2, while the remaining exhaust gas (state 10) is led into the optional cooler 3, where it is cooled if more than 55% exhaust gas is recirculated into the engine, as described above. After throttling the remaining, not recirculated exhaust gas (state 13) is finally cooled to ambient temperature (state 14/15). Again, the heat flows from all coolers could be either used as process heat or are lost to the surroundings. Both cases will be considered in the following section, with regards to exergetic efficiency. This process concept should contribute to a higher flexibility of the polygeneration process, since a wide range of exhaust gas ratios xEGR should be applicable. Output streams should be easily adjustable. In order to examine the concept, the hydrogen output and exergetic efficiencies were calculated in order to compare this process with separate conversion processes.

Results and discussion Operation parameters The chosen exhaust gas ratio effects the preheating strategy. Therefore, the temperatures of the fresh gas T2 and the

recirculated exhaust gas T12, are adjusted to the actual exhaust gas ratio xEGR. As already explained in Section “preheating temperatures and product gas composition” both temperatures are adjusted by an iterative procedure until the ignition happens within 10 crank angle after TDC. Fig. 7 shows the required temperatures (dashed lines) and equivalence ratios (solid line) as a function of xEGR. The respective temperature profiles show the effects of preheating, exhaust gas cooling and -mixing. At low exhaust gas ratios fresh gas temperature T2 and inlet temperature T3 of the engine are nearly equal at around 600 K, since the amount of exhaust gas contributing to the preheating is small and heat is mainly transferred from the exhaust gas. An increase in xEGR also results in an increasing deviation between fresh gas- and inlet temperature, until xEGR reaches 55% and the preheater is no longer needed. From this point on, the fresh gas temperature is kept constant at 298 K and the temperature of the recirculated exhaust gas T12 is reduced. The inlet and outlet temperatures T3 and T4 are slightly decreasing with increasing exhaust gas ratio. This is mainly due to dilution and the increasing content of inert gases like N2 and CO2 at high xEGR, which on one hand leads to lower outlet temperatures but also reduces the required inlet temperatures (due to the lower heat capacities). Besides process temperatures, the fresh gas equivalence ratio 4FG needs to be adjusted to the exhaust gas ratio as well. This is inevitable for a steady process, since small amounts of partially oxidized species are contained in the recirculated exhaust gas. These species are mainly remaining hydrogen, which was not separated in the membrane, carbon monoxide that was not converted in the watergas-shift reactor and unconverted fuel from the engine. In order to compensate for the oxygen lost due to the combustion of these species, the fresh gas equivalence ratio is reduced with increasing exhaust gas ratio. If the fresh gas

Please cite this article in press as: Hegner R, Atakan B, A polygeneration process concept for HCCI-engines e Modeling product gas purification and exergy losses, International Journal of Hydrogen Energy (2016), http://dx.doi.org/10.1016/j.ijhydene.2016.09.050

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Fig. 7 e Required process temperatures and fresh-gas equivalence ratios (circles, right scale) as a function of xEGR ¼ nEG/(nFG þ nEG). From T2 it can be seen whether preheating is needed.

equivalence ratio is kept constant at 4FG ¼ 2:5 the partially oxidized species in the mixture would consume most of the oxygen during combustion, resulting in an increasing amount of unconverted fuel in every subsequent engine cycle. Therefore 4FG has to be reduced from 2.5 to 0.4, as the exhaust gas recirculation ratio increases from 0 to 0.7. That way, sufficient oxygen is supplied to oxidize the fuel as well as the other species. The fresh gas equivalence ratio is also estimated together with the fresh-gas temperature T2 and the temperature of the recirculated exhaust gas T12 until the mixture ignites in the right time frame and the methane conversion is above 95%. As soon as both conditions are fulfilled, the results are taken as reasonable, leading to some non-smoothness of the results as a function of xEGR as will be seen later. A detailed compilation of results for temperature, pressure and mass flow at the different process states is shown in Table 3 for selected amounts of recirculated exhaust gas xEGR. A more detailed description of the calculation procedure of these values is given in the Supplementary material. At low xEGR high equivalence ratios can be used and a relatively high amount of hydrogen can be produced, while at high xEGR, the equivalence ratio has to be quite low, so that only a small amount of hydrogen is produced. The main product of the process shifts from hydrogen to work and heat with xEGR, which is the intended flexibility.

Exergetic efficiency The adjusted equivalence ratio and operation temperatures T2 and T12 are subsequently used to calculate the exergetic efficiencies of the process in a wide range of exhaust gas ratios, as shown in Fig. 8. Results for two cases are plotted. The upper curve marked with circles ( DT ¼ 100K) is calculated for the assumption, that the heat streams transferred from all coolers are used outside the polygeneration process with a mean temperature difference of 100 K to secondary fluids inside the coolers. The

second curve marked with squares ( Tsurr ¼ 298K) results from the assumption, that the heat is lost to the surroundings, without any downstream usage. Both curves have in common, that the efficiency is highest at low exhaust gas contents with values around 80%. With increasing xEGR this value first remains quite constant, until at xEGR ¼ 30% the efficiency starts decreasing to 60% or 40%, respectively at xEGR ¼ 70%. This is due to the adjusted equivalence ratios and the combustion of the recirculated product gas species. The combustion of recirculated hydrogen and carbon monoxide leads to higher irreversibilities and therefore causes additional exergy losses. The difference between both considered cases of heat exchange increases with increasing xEGR. At low xEGR both efficiencies are nearly equal, since most of the internal energy of the exhaust gas is used to heat up the fresh gas in the preheater. At higher xEGR the preheater is less needed and the external usage of the heat becomes crucial for an efficient process. If the heat is not used, it contributes to exergy loss and the corresponding efficiency decreases.

Process output streams From a practical point of view, the total amount of process output streams is quite relevant for downstream application and integration of the polygeneration engine in existing processes, e.g. in chemical industry. Therefore, in Fig. 9 engine power, process heat flow and hydrogen mass flow rates for one engine cylinder are shown as a function of xEGR. Especially in this diagram, but also in the previous ones, the curves are not smooth. This is a result of the iterative procedure used to determine the intake temperature and fresh gas equivalence ratio, which was explained in Section “operation parameters”. The iteration results lead to small deviations in ignition timing and fuel conversion. While the hydrogen mass flow rates are mainly unaffected by this, power and heat flow rates depend on the actual temperature- and pressure-profiles during the compression and expansion strokes. Therefore, some deviations occur and are expected, but the general trends of the profiles are still visible and will be addressed in this section. As expected, both, power output and hydrogen mass flow rates are reduced at higher exhaust gas content from originally 9.3 kW and 0.5 kg H2/h to 6.2 kW and 0.1 kg H2/h. This is a dilution effect and can also contribute to the flexibility of the polygeneration engine. By varying the amount of recirculated exhaust gas engine power and hydrogen mass flow are easily adjusted to the current demand. In contrast to the decreasing profiles of power and hydrogen, the process heat flow is even increasing the more exhaust gas is recirculated. This was already addressed in the previous section, with regards to exergetic efficiencies. At higher exhaust gas contents, the preheater becomes unnecessary and up to 8.0 kW of heat can be transferred from the exhaust gas in the various coolers. However, the heat flow decreases slightly above xEGR > 60%, because the dilution with recirculated exhaust gas results in a lower outlet temperature of the engine (see also T4 in Fig. 7). The inverse slopes of the graphs of power and heat flow rates leads to an interesting effect, where the power-to-heat ratio of the system s ¼ P=Q_ can be adjusted between 0.9 and 1.6. The power to heat value is a common value to characterize co- or polygeneration systems, consisting of power-output and heat-

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Table 3 e Operation conditions at different xEGR at different states as indicated in Fig. 6. xEGR ¼ 40%

State

1 2 3 4 5 6 7 8 9 10 11 12 13 14 15

xEGR ¼ 55%

xEGR ¼ 70%

T/K

P/kPa

_ m/kg/h

T/K

P/kPa

_ m/kg/h

T/K

P/kPa

_ m/kg/h

298.0 490.0 586.0 868.4 743.4 723.0 723.0 723.0 298.2 723.0 723.0 723.0 723.0 529.1 298.0

101.33 101.33 101.33 165.68 165.68 165.68 165.68 101.33 101.33 165.68 165.68 101.33 101.33 101.33 101.33

12.80 12.80 22.02 22.02 22.02 22.20 22.20 0.26 0.26 21.94 21.94 9.22 12.73 12.73 12.73

298.0 298.0 543.4 843.0 723.0 723.0 723.0 723.0 298.2 723.0 723.0 723.0 723.0 723.0 298.0

101.33 101.33 101.33 165.84 165.84 165.84 165.84 101.33 101.33 165.84 165.84 101.33 101.33 101.33 101.33

10.60 10.60 24.26 24.26 24.26 24.26 24.26 0.16 0.16 24.10 24.10 13.66 10.44 10.44 10.44

298.0 298.0 494.7 770.9 723.0 723.0 723.0 723.0 298.2 723.0 570.0 570.0 570.0 570.0 298.0

101.33 101.33 101.33 161.34 161.34 161.34 161.34 101.33 101.33 161.34 161.34 101.33 101.33 101.33 101.33

8.02 8.02 27.12 27.12 27.12 27.12 27.12 0.09 0.09 27.02 27.02 19.10 7.92 7.92 7.92

Fig. 8 e Exergetic efficiencies as function of exhaust gas content xEGR ¼ nEG/(nFG þ nEG) with different assumptions regarding process heat usage, as explained in the text.

output. Usually this ratio is restricted to a narrow range, dependent on the system itself, e.g. around 1.1 for IC-engines and 0.6 for gas turbine co-generation [30]. The comparatively wide range for the proposed engine polygeneration system makes it suitable for various applications from chemical industry to food processing and textile fabrication [31].

Comparison with single-product systems Lastly the possible advantages of the polygeneration approach shall be discussed by comparing it with the separated generation of power, heat and hydrogen. For comparison, power and heat shall be supplied by a conventional HCCI-engine, operating in the fuel lean regime at equivalence ratios between 0.3 and 0.5, in contrast to the fuel rich polygeneration approach. HCCI engines are known to have high efficiencies, therefore a fair comparison should be possible. The respective

Fig. 9 e Power-, heat-flow rates (both left scale)and hydrogen-mass flow rates (squares, right scale) as output of the polygeneration process dependent on xEGR ¼ nEG/ (nFG þ nEG).

calculations were performed using the HCCI-simulation model described in Section “modeling” and the inlet temperatures in Fig. 2. For this comparison, hydrogen is assumed to be produced by a quite efficient steam methane reforming (SMR) process. In this process, hydrogen is produced catalytically by means of steam methane reforming (see Eq. (13)) and subsequent water-gas-shift reaction (Eq. (12)). The hydrogen is also separated via metallic membranes. CH4 þ H2 O #CO þ 3H2

(13)

Quantitative values were taken from a publication of Simpson and Lutz [19]. Table 4 summarizes important key values of the polygeneration process and the compared processes. The results indicate, that specific power-, heat- and hydrogen production of the polygeneration process are lower, compared to the conventional HCCI-engine or steam

Please cite this article in press as: Hegner R, Atakan B, A polygeneration process concept for HCCI-engines e Modeling product gas purification and exergy losses, International Journal of Hydrogen Energy (2016), http://dx.doi.org/10.1016/j.ijhydene.2016.09.050

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Fig. 10 e Fuel consumption of the polygeneration process and cumulative fuel consumption of the comparative processes (fuel lean HCCI engine þ steam methane reforming process).

theoretically. The purpose of the polygenerating engine was hydrogen co-production via partial oxidation of the fuel methane. An HCCI-single-zone model was used to simulate the in-cylinder process and to obtain realistic product gas compositions. Since the product gas contained several byproducts, the hydrogen separation process was included in the process concept. Another challenge addressed during the development of the process was the fresh gas preheating, which was inevitable due to the high auto-ignition temperatures of methane. This was realized by a combined use of heat exchangers and recirculation of a part of the hot exhaust gas. The results revealed that the recirculated exhaust gas also carried small amounts of partially oxidized species like hydrogen and carbon monoxide into the subsequent engine cycle. Since the combustion of these species would result in lower conversion of the actual fuel, the fuel-air equivalence ratio was adjusted to the amount of recirculated exhaust gas. Although this led to a decrease in exergetic efficiency at higher exhaust gas contents, a maximal efficiency of 80% is still considerably higher than the efficiencies of conventional processes for hydrogen production, like steam methane reforming [19]. However, the usage of the process heat in a possible downstream applications or alternatively, its loss to the surrounding has an high impact on the exergetic efficiency. Besides its contribution to preheating, the main advantage of exhaust gas recirculation is the easy adjustment

Table 4 e Comparison of the polygeneration process with HCCI-engine and SMR.

Output 4 Spec. Output

Ex. efficiency

Polygeneration

HCCI-engine

Steam methane reforming

Power þ Heat þ H2 2.5 Work: 8e23 MJ/kgCH4 Heat: 5e27 MJ/kgCH4 H2: 0.2 kg/kgCH4 58-80%

Power þ Heat 0.3e0.5 Work: 24 MJ/kgCH4 Heat: 7 MJ/kgCH4

H2 1 H2: 0.28 kg/kgCH4

58%

63%

reforming. This is expected, since the polygeneration process is designed to have a low power output in favor of the coproduction of hydrogen. However, the exergetic efficiency of the polygeneration process surpasses the separate systems by up to 20%, indicating that energy conversion in polygeneration is improved and less fuel is consumed. To confirm this, the fuel consumption of the polygeneration process and the summed up fuel consumption of a conventional HCCI-engine and a steam methane reforming process, producing the same power and hydrogen flow rate are compared in Fig. 10. The results demonstrate, that the engine polygeneration process can significantly reduce fuel consumption in fuel conversion processes. Especially at very low and very high exhaust gas ratios the fuel savings are remarkable, with 20% lower fuel consumption at xEGR ¼ 5% and 40% at xEGR ¼ 70%.

of the engine power, heat and hydrogen output, by varying the amount of recirculated exhaust gas. Lastly a comparison of the polygeneration approach with the separated production of power, heat and hydrogen demonstrated, that up to 40% fuel savings could be accomplished by polygeneration. All results indicate, that the production of hydrogen in piston engine polygeneration processes, has several advantages, compared to conventional hydrogen production. It includes the possibility to cogenerate work and heat as well, with a large flexibility towards the desired values. This all is accomplished with low exergy losses, due to the thermodynamically favorable conversion.

Acknowledgement Conclusions A polygeneration process concept which is based on a fuelrich operated HCCI-engine was developed and investigated

Financial support of this work (AT24/13-1) by the Deutsche Forschungsgemeinschaft within the framework of the DFG research unit FOR 1993 ‘Multi-functional conversion of chemical species and energy’ is gratefully acknowledged.

Please cite this article in press as: Hegner R, Atakan B, A polygeneration process concept for HCCI-engines e Modeling product gas purification and exergy losses, International Journal of Hydrogen Energy (2016), http://dx.doi.org/10.1016/j.ijhydene.2016.09.050

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Appendix A. Supplementary data Supplementary data related to this article can be found at http://dx.doi.org/10.1016/j.ijhydene.2016.09.050.

references

[1] Mohamad TI, How HG. Part-load performance and emissions of a spark ignition engine fueled with RON95 and RON97 gasoline. Technical viewpoint on Malaysia’s fuel price debate. Energy Convers Manag 2014;88:928e35. ~ oz de Escalona JM, Sa  nchez D, Chacartegui R, Sa  nchez T. [2] Mun Part-load analysis of gas turbine & ORC combined cycles. Appl Therm Eng 2012;36:63e72. [3] Chicco G, Mancarella P. Distributed multi-generation. A comprehensive view. Renew Sustain Energy Rev 2009;13:535e51. [4] Wang Z, Zheng D, Jin H. A novel polygeneration system integrating the acetylene production process and fuel cell. Int J Hydrogen Energy 2007;32:4030e9. [5] He C, Feng X. Evaluation indicators for energy-chemical systems with multi-feed and multi-product. Energy 2012;43:344e54. [6] Spencer JD, Moton JM, Gibbons WT, Gluesenkamp K, Ahmed II, Taverner AM, et al. Design of a combined heat, hydrogen, and power plant from university campus waste streams. Design of a combined heat, hydrogen, and power plant from university campus waste streams. Int J Hydrogen Energy 2013;38:4889e900. [7] Wang Z, Zheng D, Jin H. Energy integration of acetylene and power polygeneration by flowrate-exergy diagram. Appl Energy 2009;86:372e9. [8] Margalef P, Brown T, Brouwer J, Samuelsen S. Conceptual design and configuration performance analyses of polygenerating high temperature fuel cells. Int J Hydrogen Energy 2011;36:10044e56. [9] Cormos C-C. Renewable hydrogen production concepts from bioethanol reforming with carbon capture. Int J Hydrogen Energy 2014;39:5597e606. [10] Yu Z, Han J, Cao X. Investigation on performance of an integrated solid oxide fuel cell and absorption chiller trigeneration system. Int J Hydrogen Energy 2011;36:12561e73. [11] Karim GA, Wierzba I. The production of hydrogen through the uncatalyzed partial oxidation of methane in an internal combustion engine. Int J Hydrogen Energy 2008;33:2105e10. [12] von Szeszich L. Herstellung von Synthesegas im Otto-Motor bei gleichzeitiger Arbeitsgewinnung. Chem Ing Techn 1956;28:190e5. [13] Yang YC, Lim MS, Chun YN. The syngas production by partial oxidation using a homogeneous charge compression ignition engine. Fuel Process Technol 2009;90:553e7. [14] Morsy MH. Modeling study on the production of hydrogen/ syngas via partial oxidation using a homogeneous charge compression ignition engine fueled with natural gas. Int J Hydrogen Energy 2014;39:1096e104.

11

[15] Mcmillian M, Lawson S. Experimental and modeling study of hydrogen/syngas production and particulate emissions from a natural gas-fueled partial oxidation engine. Int J Hydrogen Energy 2006;31:847e60. [16] Lim EG, Dames EE, Cedrone KD, Acocella AJ, Needham TR, Arce A, et al. The engine reformer. Syngas production in an engine for compact gas-to-liquids synthesis. The engine reformer. Syngas production in an engine for compact gasto-liquids synthesis. Can J Chem Eng 2016;94:623e35. [17] Gossler H, Deutschmann O. Numerical optimization and reaction flow analysis of syngas production via partial oxidation of natural gas in internal combustion engines. Int J Hydrogen Energy 2015;40:11046e58. [18] Sonntag RE, Borgnakke C, van Wylen GJ. Fundamentals of thermodynamics. 6th ed. New York: Wiley; 2003. [19] Simpson A, Lutz A. Exergy analysis of hydrogen production via steam methane reforming. Int J Hydrogen Energy 2007;32:4811e20. [20] Goodwin DG, Moffat HK, Speth RL. Cantera: an objectoriented software toolkit for chemical kinetics, thermodynamics, and transport processes.: Version 2.7.1. 2014. http://www.cantera.org. [21] Wang Hai, You Xiaoqing, Joshi Ameya V, Davis Scott G, Laskin Alexander, Egolfopoulos Fokion, et al. USC Mech Version II. High-temperature combustion reaction model of H2/CO/C1-C4 compounds. 2007. http://ignis.usc.edu/USC_ Mech_II.htm. [22] Zheng J, Caton JA. Use of a single-zone thermodynamic model with detailed chemistry to study a natural gas fueled homogeneous charge compression ignition engine. Energy Convers Manag 2012;53:298e304. [23] Heywood JB. Internal combustion engine fundamentals. Singapore: Mc Graw Hill International Editions; 1988. [24] Woschni G. A Universally Applicable Equation for the Instantaneous Heat Transfer Coefficient in the Internal Combustion Engine. SAE International 1967. [25] Kotas TJ. The exergy method of thermal plant analysis. 2012nd ed. Malabar Fla: Krieger; 2012. [26] Merker G. Verbrennungsmotoren: Simulation der Verbrennung und Schadstoffbildung. Mit 15 Tabellen. 3rd ed. Wiesbaden: Teubner; 2006. [27] Sanz R, Calles JA, Alique D, Furones L. H2 production via water gas shift in a composite Pd membrane reactor prepared by the pore-plating method. Int J Hydrogen Energy 2014;39:4739e48. [28] Adhikari S, Fernando S. Hydrogen membrane separation techniques. Ind Eng Chem Res 2006;45:875e81. [29] Hajjaji N, Pons M-N, Houas A, Renaudin V. Exergy analysis. An efficient tool for understanding and improving hydrogen production via the steam methane reforming process. Energy Policy 2012;42:392e9. [30] Frangopoulos CA. A method to determine the power to heat ratio, the cogenerated electricity and the primary energy savings of cogeneration systems after the European Directive. Energy 2012;45:52e61. [31] Petchers N. Combined heating, cooling & power handbook: technologies & applications an integrated approach to energy resource optimization. 2nd ed. Lilburn, GA, New York, NY: The Fairmont Press; 2012.

Please cite this article in press as: Hegner R, Atakan B, A polygeneration process concept for HCCI-engines e Modeling product gas purification and exergy losses, International Journal of Hydrogen Energy (2016), http://dx.doi.org/10.1016/j.ijhydene.2016.09.050