1. Introduction
Renewable energy production technologies are going to play a significant role in the immediate future due to global warming and its significant consequences, Therefore, it is essential to find new, effective solutions that allow for the integration of sustainable energy production techniques into the current existing systems thereby decreasing the emissions. In order to use the most energy of the renewable sources, it will be essential to use such solutions with polygeneration system to maximize the energy use. Examples such polygeneration can be the production of electricity, fuel and fresh water (instead of dissipating the excess heat to the environment).
Using electrolysis technology such as solid oxide electrolyte cells (SOECs) is one way to store the excess energy from renewable sources when the production is higher than the demand. One can then use the stored fuel to generate power and produce other useful outputs such as heat, cooling and fresh water through solid oxide fuel cells (SOFCs) when the renewable source is low, like on a calm day when using wind energy or during nighttime when using sun energy. This indicates that, then one can use a reversible solid oxide cell (RSOC) to produce synthetic fuel from electricity, or to produce electricity from fuel when reversed.
Several studies on SOEC systems have been conducted; for example, [
1] reviewed technological developments in hydrogen production from an SOEC system in terms of materials, cell configuration designs, electrode depolarizations and mathematical modelling. Reference [
2] presented a newly designed apparatus for testing single solid oxide cells in both fuel cell and electrolysis modes, which in turn showed performance improvements when running in electrolysis mode. Reference [
3] conducted energy and exergy studies of a solid oxide cell plant to evaluate system performance in terms of energy losses, exergy destruction, and hydrogen production efficiency. Reference [
4] experimented on a 16 cm
2 planar solid oxide cell to test it as a fuel cell and electrolyser at various pressures ranging from 0.4 bar to 10 bar and found out that the internal resistance decreases with increasing pressure. Reference [
5] showed the successful reversible operation of a dual-mode cell through electrochemical tests carried out by impedance spectroscopy and thereby proving its feasibility of the concept. Reference [
6] demonstrated the heat spreading capabilities and power limitations of high-temperature applications in SOEC/SOFC stacks through an experimental study.
The European Union (EU) energy policy includes EU low-carbon roadmap milestone, which indicates a 40% reduction in carbon emissions and 30% EU-wide share of renewables. The increased renewable energy sources penetration and their irregular energy production have led to the emerging the need for energy storage technologies. This in turn means that EU energy market is changing towards application of Power to Gas (P2G) systems. Many studies discuss P2G systems based on water electrolysis, and this study mentions some of these studies. Reference [
7] studied the techno-economics of P2G systems and concluded that these systems require subsides to overcome economic barriers. The study of [
8] showed that an economic dispatch model involving wind power and a battery-energy storage system integrated with P2G helps to reduce the amount of wind power curtailment and stabilize fluctuations in the wind power output. Reference [
9] performed a techno-economic analysis of eight different pathways scenarios including hydrogen via electrolysis. Reference [
10] highlighted P2G processes and their products as a promising solution to the problems faced by the transport sector, while [
11] discussed the role of and potential of deploying P2G applications for the generation, transmission and distribution of electricity. Reference [
12] suggested that a smart energy network should also include a P2G system via e.g. water electrolysis, while [
13] argued that the most economically affordable technology for hydrogen production is the electrolysis system using surplus wind energy in Ireland. Reference [
14] presented a scenario for integrating P2G with a multi-source microgrid including wind energy, micro turbine, etc. Reference [
15] performed a study on the techno-economic and environmental assessment of bio-methane production via biogas upgrading and P2G technology. Reference [
16] uses a different approach, which is integrating gasification with P2G through water electrolysis. However, none of these studies present a detail plant balance as in the present study.
Direct contact membrane distillation (DCMD) is a thermal separation process that allows only water vapor (or other volatiles) to pass through a micro-porous hydrophobic membrane while no impurities such as salt can cross it. The reason is the driving process is created by the vapor pressure gradient (and by temperature difference) over both sides of the membrane. Reference [
17] reviewed the desalination of seawater using DCMD systems, and their performance from laboratory scale to pilot projects thereby proving the feasibility of concept. Reference [
18] showed by experiments that hot water at 80 °C under optimum conditions and with an optimum membrane selection can separate 99.99% of salt from the water. Desalination systems powered by waste heat are an attractive solution that can address the worldwide water-shortage problem, without contributing significant greenhouse gas emissions. Reference [
19] also stated that low-temperature DCMD systems are a promising system for freshwater production. The study by [
20] indicates that the experimental data agrees very well with the calculated results in terms of vapor mass flux and overall heat transfer coefficients. In addition, such a technique has a great advantage because it works at even lower temperatures, such as 40 °C, suggesting the use of lower temperature sources and thereby avoiding the latent heat of water [
21].
This study presents a novel polygeneration system (power-to-gas) that uses the excess energy from the wind turbines to feed a RSOC system thereby producing fuel for later use. Further, the system design recovers the waste heat from the RSOC system to produce heating, cooling and fresh water in addition to fuel. Such a system will results in a flexible polygeneration plant driven mainly by wind energy that can regulate different output combinations of hydrogen, heat, cool and fresh water. The study first presents a complete balance of the plant and then offers alternative system designs for different demands. Further, this study analyses the performance of each design thermodynamically. To the best of the author’s knowledge, no similar studies exist in the open literature. The objective of this study is to discover alternative possible future power-to-gas technologies for integration into existing systems, which might be of interest for future power generation needs.
3. Proposed Plant Schemes
Figure 3 presents the proposed plant scheme in this study. As shown, PTSC preheats the water to 350 °C (node 3) before entering the anode preheater (AP). In other words, it acts as a steam generator for the RSOC and this is the main reason why this study uses PTSC. As mentioned above the term anode refers to the fuel cell mode. The water (now steam) is preheated in the anode preheater to about 630 °C before entering the RSOC. The temperature of the off-fuel (node 6) is 750 °C, which is used to preheat the steam in the anode preheater. The off-fuel after the anode preheater (node 7) is then first cooled down in a district heating heat exchanger (DH2) and then is sent to a condenser for separating the H
2 and H
2O. It shall be noted that the off-fuel after the RSOC is a mixture of H
2 and H
2O. This mixture depends on the utilization factor of the RSOC. The higher the utilization factor is the lower the amount of water in the mixture will be. The hydrogen and water mixture (off-fuel) after the RSOC is separated at 100 °C.
The design extracts some of the steam generated by the PTSC for the district heating (heat exchanger denoted as DH1 in the figure). The reason for this extraction is to regulate the temperature of the steam entering the RSOC. This is another reason why the proposed design uses a solar collector to generate steam. Later on, the analysis shows that the amount of this extraction is very important when wind velocity varies. The off-oxygen after the RSOC, which is separated from the steam in the RSOC (node 23), has a temperature of about 750 °C. One can use such hot stream for different purposes such as generating heat for the district heating (DH3) network, generating cooling for the district cooling network, or producing fresh water from a distillation unit (DCMD in this study). Thus, such designs offers a variety of possibilities depending on the location where the plant is to be installed. The operating temperature of DCMD is assumed to be 80 °C allowing 10 °C for the pinch temperature then off-oxygen is cooled down to 90 °C.
The driving forces of the RSOC is the excess energy from the wind turbines. In Denmark, the installed wind turbines are getting larger and larger and therefore in many hours of the year the electricity produced by the wind energy exceeds the demand. The suggested plant uses this excess electricity to generate fuel and at the same time, some other useful production. The study assumes a thermo-neutral voltage for all simulations carried out. This means that depending on the power supplied to the electrolyser, the cell voltage varies so that there is no need to supply heat to the electrolyser. Note that if the cell voltage is lower than the thermo-neutral voltage then the electrolyser needs heat at a very high temperature, which is not feasible.
The supply temperature for the district heating is 100 °C while its return temperature is 50 °C. These values are based on the current technology in Denmark. New generation district heating under development will have supply temperature at about 50 °C to 60 °C.
An alternative plant design replaces the DH3 in the off-oxygen stream with an absorption chiller, as shown in
Figure 4. This plant is able to produce cooling (in addition to the heating) whenever cooling is needed, e.g. during summertime if located in a colder region, or year around if located in a warm region. Note that this particular design uses DH1 and DH2 for domestic hot water production (showering, washing, etc.). Therefore, such a combination provides many opportunities depending on the location of the plant.
A third alternative is proposed in
Figure 5, wherein the absorption chiller is replaced with a DCMD unit to produce fresh water. Fresh water is becoming scarce in many areas and the need for such units is becoming more and more important, and therefore it is studied here.
The efficiency defined above (Equation (9)) does not take account the heat production, cool production and fresh water production. It only defines fuel production. Therefore, there is a need to define a new efficiency, which accounts for other production besides the fuel production (
Qprod). Thus, the following efficiency is defined:
Such a definition calls for plant energy efficiency or plant utilization efficiency. Another point to be mentioned is that the solar energy is free and therefore one can assume that its contribution to the efficiency can be neglected. Therefore, the following efficiencies can be defined:
Obviously, the plant efficiency according to Equation (20) may be larger than unity under certain circumstances and the reason is that it neglects the free heat input from the solar energy to the system. In this study
Qprod is then the summation of heat production, cool production and freshwater production as:
4. Results and Discussions
Figure 6 presents wind turbine performance curves. It shows that for any design there exists a wind velocity for which the power output is maximum (
Figure 6a). It also demonstrates that for each design and at a constant wind velocity there exists a rotational speed for which turbine power is maximum (
Figure 6b).
The figure also indicates that wind power decreases when the wind velocity is above 12 m/s (the default value for the present design). Another conclusion is that for each pitch angle there is a rotational speed at which the power is maximized. This indicates that to operate the wind turbines at their peak efficiency one needs to design a variable shaft speed, but at the expense of additional cost. It should be noted that for safety reasons, most companies also design the wind power in a way that when the power reaches its maximum then the rational speed is kept constant and does not respond to additional wind speed increases. Such a design avoids mechanical fracture, damage and failure of the blades. However, this study does not take this into account for structural modelling. The wind turbines does not respond to wind powers below about 5–6 m/s as shown in
Figure 6a.
Figure 7 shows the optimum power of a wind turbine for different wind velocities and attack angles. As seen, the power is maximized at a wind velocity of 10 m/s and an attack angle of about 30°. Therefore, hereafter all calculations consider these values.
4.1. Plant with District Heating Only
As demonstrated above, wind velocity is an important parameter for investigation, due to the power coefficient of the wind turbines. Wind turbines’ electrical power strongly depends on the wind power (wind speed) which directly affects the hydrogen production through the electrolyser system.
Figure 8 presents the SOEC performance when the wind speed is changed.
As demonstrated above, the power of the turbine first increases to maximize at about 10 m/s (wind velocity) and then starts to decrease at wind velocities higher than about 11 m/s and therefore the power feed to the electrolyser is maximized at a wind velocity of about 10 m/s (parabolic shape). This results in a parabolic shape of the current densities as well as cell voltage. The current density and cell voltage increase when wind velocity increases from 6 m/s to 10 m/s. On the other hand, the current density and cell voltage decrease when the wind velocity is above 10 m/s. The current density and cell voltage reach 1208 mA/cm2 and 1.366 V, respectively, when the wind velocity is 10 m/s. We note also that this study takes into account the thermo-neutral voltage (no heat supplied to the electrolyser) and calculates this voltage. Since power supplied to the electrolyser decreases at wind velocities above 10 m/s, then hydrogen production first increases from 945 kg/day to maximize to 2197 kg/day and then starts to decrease and reaches to 1287 kg/day when the wind velocity is 16 m/s.
Figure 9 displays heat production as well as heat and power consumptions by the system with district heating only (c.f.
Figure 3). Note that plant heat consumption is due to the solar energy through the PTSC while the plant power consumption is coming from wind turbines. Heat consumption through the PTSC (from solar energy) is constant since the number of PTSCs does not change and the temperature out of the PTSC is constant at 350 °C. The plant power consumption varies with the wind velocity and maximizes at 10 m/s which is about 4115 kW.
Plant total heat production (for district heating) shows a similar pattern (behaviour) as the power supplied to the electrolyser. Heat production maximizes at a wind velocity of 10 m/s, which is 865 kJ/s. Such a pattern of behaviour of the heat production is mainly due to the heat production of DH2, which is located at the off-hydrogen side and depends strongly on the electrolyser performance in terms of hydrogen production. Thus, heat production from DH2 located at the off-fuel side of the electrolyser increases/decreases as a direct consequence when the electricity supplied to the SOEC increases/decreases. Thus, heat production by DH2 is maximized when the wind velocity is 10 m/s (601 kJ/s). The reason is that steam production (mass flow) of the PTSC is constant (constant size and solar radiation) and therefore more steam goes through electrolysis as power to the electrolysis increases. This in turns means that less steam goes through the splitter just after the PTSC. Lower steam through this splitter causes less heat production through the DH1, which happens when the wind velocity increases from 6 m/s to 10 m/s, meaning that there exists less excess of steam for DH1. In other words, DH1 produces more heat as the electrolyser performance decreases. It is now obvious why the design includes a splitter after the PTSC. Further, heat production by DH1 is minimized when the wind velocity is 10 m/s (132 kJ/s). Heat production by DH3 (located on the off-oxygen side) strongly depends on the electrolysis performance, as better performance means more mass flow of off-oxygen. It is maximized when the wind velocity is 10 m/s (132 kJ/s).
Figure 10 exhibits the electrolyser system efficiency as well as the plant efficiency when heating production is included. Plant performance in terms of efficiencies maximizes when the wind velocity is 10 m/s (Equations (9) and (18)). Electrolysis efficiency (Equation (9)) or hydrogen production efficiency increases from 24% to 39% when the wind velocity increases from 6 m/s to 10 m/s. Increasing the wind velocity from 10 m/s to 16 m/s causes the hydrogen efficiency to decrease from 39% to 29%. Similarly, the plant efficiency (Equation (18)) maximizes when the wind velocity is 10 m/s to reach a value of about 50%. On the other hand, neglecting the heat input by the solar energy (free heat) causes the plant performance to have a minimum value when the wind velocity is 10 m/s. Electrolyser efficiency (Equation (19)) or hydrogen production efficiency decreases from 76% to about 74% when the wind velocity increases from 6 m/s to 10 m/s. Further increases in wind velocity increase the hydrogen production efficiency back to 76% when the wind velocity reaches 16 m/s. However, such a variation is small and negligible. The minimum plant energy efficiency (Equation (20)) is about 95%, which happens when the wind velocity is 10 m/s. These results are encouraging and demonstrate the importance of including renewable energy systems in current energy systems.
As mentioned above, the supply temperature to the district heating network is 100 °C (current and most used technology). This indicates that some energy is lost from the system without being recovered. Decreasing the DH supply temperature to 50 °C (future DH generation under development) decreases the energy dissipation to the environment and thereby increases the plant efficiency. Note that Qprod in Equations (18) and (20) accounts for the summation of all heat generated for district heating (DH1, DH2 plus DH3).
As noted, the efficiency according to Equation (20) can be larger than 100% because this equation neglects the heat added to system from the solar energy (which varies from 95% to 115%). Including this free heat, the plant efficiency varies from 36% to about 50%, depending on the wind velocity.
4.2. Plant with District Heating and Cooling
The results for the plant with both DH and an absorption chiller are revealed in
Figure 11.
Off-air after the desorber is dissipated at 90 °C, well above the dew point. Now, Qprod in Equations (18) and (20) accounts for the summation of heat and cool generation. Multiplying the mass flow of the cooling flow with the enthalpy difference over the evaporator, we can calculate the generated cooling effect. The results obtained here are very similar to the previous case. SOEC hydrogen production and plant efficiencies (Equations (9) and (18)) are maximized when the wind speed is 10 m/s while neglecting the free energy from the Sun then these efficiencies (Equations (19) and (20)) is minimized when the wind speed is 10 m/s.
4.3. Plant with District Heating and Freshwater
Figure 12 demonstrates the results obtained here from a plant, which generates both heat and fresh water in addition to producing hydrogen (c.f.
Figure 5). Again, the results are similar as for the previous case, signifying that the definition in Equations (18) and (20) can be used for all polygeneration systems.
In this case,
Qprod in Equations (18) and (20) is the summation of heat and fresh water generation. Since the fresh water production (kg/s) does not have the same dimensions as the heat (J/s) then this study defines the fresh water generation (J/s) as the energy difference between the inlet and outlet of the DCMD, which is the energy difference between the produced fresh water out of the DCMD and input water to the DCMD,
Similarly for seawater one would have:
and finally the performance of DCMD can be defined as:
Finally, the performance of the absorption chiller is calculated to be 0.617, which is the ratio between the heat released in the evaporator (cooling effect) over the heat absorbed in the desorber. The performance of the DCMD is calculated to be 0.934, which is the ratio between the heat absorbed by the fresh water, and the heat lost by the seawater in the DCMD (defined above).
As noted above, wind velocity has a significant effect on the plant performance. On the other hand, the solar radiation changes significantly during a day and therefore, it may have some effect on the plant performance. Therefore, a dynamic model may better capture plant performance by knowing solar radiation and wind velocity hour per hour for a specified location. Such data are usually available from the weather data for the region where the plant is to be placed.
4.4. Effect of Solar Radiation
Since the plant utilizes a PTSC to generate steam then the effect of solar radiation on hydrogen fresh water production is also studied, see
Figure 13. As documented, increasing the solar input energy increases fresh water production while hydrogen production remains almost unchanged. The reason is that wind power is constant and therefore power input to the electrolyser does not change. Consequently, the hydrogen production remains also unchanged while due to the higher solar radiation more steam will be generated which bypasses the separator before the cathode preheater (c.f.
Figure 3). Note that in this case the DH in
Figure 3 are replaced with DCMDs.