- Email: [email protected]
Cathode–anode side interaction in SOFC hybrid systems
Applied Energy 105 (2013) 369–379
Contents lists available at SciVerse ScienceDirect
Applied Energy journal homepage: www.elsevier.com/locate/apenergy
Cathode–anode side interaction in SOFC hybrid systems Mario L. Ferrari ⇑, Aristide F. Massardo Thermochemical Power Group (TPG), DIME, University of Genoa, Italy
h i g h l i g h t s " Cathode–anode interaction issues. " Hybrid system emulator with anodic recirculation. " Grid-connected and stand-alone tests. " Cathode–anode differential pressure. " Control system aspects.
a r t i c l e
i n f o
Article history: Received 31 July 2012 Received in revised form 3 January 2013 Accepted 8 January 2013 Available online 9 February 2013 Keywords: SOFC hybrid system Microturbine Test rig Cathode–anode interaction
a b s t r a c t Cathode–anode interaction, mainly based on cathode versus anode volume influence, recirculation performance, and turbomachinery integration, is an important issue for pressurised SOFC hybrid systems, and this aspect must be carefully considered to prevent fuel cell ceramic material failures through a reliable control system. Over the last 10 years, several theoretical analyses of this issue have been carried out at the University of Genoa. These interaction studies have been analysed and an experimental approach (for model validation, system development and prototype design activities) has been applied using emulator facilities or real plants. In particular, general hybrid system layouts based on the coupling of pressurized SOFC stacks of different geometries (planar, tubular, etc.) with a gas turbine bottoming cycle have been investigated using the hybrid system emulator facility of the University of Genoa. The experimental results are focused on the interaction between gas turbine and anodic circuit and on cathode–anode differential pressure behaviour for design, off-design and transient hybrid system operative conditions. The information obtained in these tests is essential to understand the main features of the variables that drive the phenomena and to design a suitable control system that can mitigate the differential pressure values during all plant operating conditions. Ó 2013 Elsevier Ltd. All rights reserved.
1. Introduction The increase in demand for electrical energy [1], coupled with a heightened awareness of environmental pollution problems [2] related to power system emissions, provides a strong incentive for researchers to develop more efficient plants for power generation. In addition, since conventional fossil fuel generation activity is the major source of greenhouse gas emissions [3], this field of industry certainly needs to improve to meet international targets [4]. In this scenario, high temperature fuel cell hybrid systems are a possible effective solution to environmental and efficiency issues [5–7]. This technology is based on the coupling of a fuel cell stack (able to convert the chemical energy of fuel directly into electricity without the efficiency limitations typical of thermodynamic cycles) ⇑ Corresponding author. E-mail addresses: [email protected] (M.L. Ferrari), [email protected] (A.F. Massardo). 0306-2619/$ - see front matter Ó 2013 Elsevier Ltd. All rights reserved. http://dx.doi.org/10.1016/j.apenergy.2013.01.029
with a conventional gas turbine. In addition to efficiency benefits, hybrid systems produce low pollutant [8] exhaust gases at temperature values useful for co-generative applications [8–10]. Two pressurised hybrid system general layouts (Fig. 1), based on solid oxide fuel cell (SOFC) technology developed by Siemens Westinghouse [5,11,12] and Rolls-Royce Fuel Cell Systems [13,14], are considered in this paper. The fuel cell, usually located between compressor and turbine (pressurised solution), is equipped with an ejector based anodic recirculation and an offgas burner. Fuel is fed into the primary duct of an anodic ejector [15] (see Fig. 2 for ejector nomenclature details). A reforming section may be present upstream of the anode inlet [11] or located internally in the fuel cell [14]. In the ‘‘Layout A’’ a standard recuperator is included to pre-heat the air flow upstream of the fuel cell inlet exploiting turbine exhaust thermal energy [15]. This configuration is an interesting and efficient solution for an SOFC stack equipped with an air pre-heating heat exchanger (or tube), such as the tubular fuel cells developed by Siemens Westinghouse
370
M.L. Ferrari, A.F. Massardo / Applied Energy 105 (2013) 369–379
Nomenclature F N TOT V
Acronyms AC alternate current DC direct current Ex heat exchanger mGT micro-gas turbine OGB off-gas burner REC recuperator SOFC solid oxide fuel cell TPG thermochemical power group UDP user datagram protocol WHEx water heat exchanger (for co-generation)
recirculation ratio () rotational speed (rpm) turbine outlet temperature (K) volume (m3)
Greek symbols e effectiveness Subscripts 0 on-design an anodic cat cathodic in inlet out outlet
Variables g 9.81 (m/s2)
[5,11]. ‘‘Layout B’’ is based on a cathodic recirculation loop carried out through a cathodic ejector [15]. This recirculation layout is less efficient for air heating but useful to avoid recuperator costs and high temperature heat exchanger constraints (and also very interesting for the thermal control of the stack [14,16]). Cathode and anode sides volume sizes are different for the two schemes shown in Fig. 1. This is related to the different fuel cell technology: while the tubular cell of the ‘‘Layout A’’ has an anodic volume significantly higher than the cathodic one (cathode/anode volume ratio around 0.5 [17]), the reverse (cathode/anode volume ratio around 4 [16]) applies to fuel cells not equipped with high temperature air pre-heating systems (e.g. planar SOFCs). Moreover, the two layouts have different values for pressure loss in the cell sides. For this reason, even if cathode and anode flows are mixed in the OGB in both layouts, different pressure values are possible in the upstream ducts between cathode and anode sides, especially during hybrid system part-load and time-dependent operations. Thermo-chemical aspects in time-dependent modes (e.g. startup, shutdown and load change phases) and control performance
depend on dynamic constants related to volume dimension (besides being influenced by ejector and machine performance, which are essential for the main influence properties: ejector recirculation ratio, ejector pressure increase, compressor delivery pressure). Cathode–anode interaction analysis (mainly the results of the pressure difference between fuel cell sides) is important to understand the effects on the bottoming cycle gas turbine components and for SOFC material safe management. In details, anodic side performance affects the cathode side due to the additional mass flow rate from fuel injection and the related modification of energy input. This generates possibly significant variations in cathodic side properties, such as machine rotational speed (or turbine outlet temperature), compressor outlet pressure and recuperator performance (see ‘‘Layout A’’ in Fig. 1) or cathodic recirculation aspects (‘‘Layout B’’ in Fig. 1). Since this behaviour can produce dangerous effects (e.g. surge events for compressor outlet pressure increase) at off-design conditions (or during transient operations), a complete analysis is essential.
Fig. 1. Hybrid systems: general layouts.
M.L. Ferrari, A.F. Massardo / Applied Energy 105 (2013) 369–379
Fig. 2. Ejector details.
Cell cathode–anode differential pressure behaviour is a very important aspect of this interaction. Apart from some patents [18,19] based on cathode–anode differential pressure management for low temperature fuel cell systems (turbomachinery effects are quite limited), no detailed theoretical investigations and experimental data have been published on this topic for SOFC hybrid systems. Moreover, no experimental data have been published on cathode–anode differential pressure trends at part-load and time-dependent conditions. On the other hand, many authors have been involved in theoretical activities analysing whole hybrid systems at different operative conditions [14,17,20–22], including control system issues [17]. For instance, cathode–anode interaction aspects were analysed in a previous model-based activity [23], in which the Authors calculated a 15–20 mbar cathode–anode pressure increase for each 10% power decrease (from nominal load value) if no control system is included for air flow. A subsequent publication [17] demonstrated how this mechanical stress could be avoided by controlling the gas turbine rotational speed. Even if the cathode–anode stress problem can be studied at stack level to increase ceramic material resistance performance, system level studies are needed to analyse the stress trend during part-load and time-dependent conditions. Since this kind of data is not available in open literature (manufacturers do not usually provide details on these aspects), a broad experimental support is mandatory to completely understand the phenomenon, validate models, and undertake prototype development activities. To generate experimental data representative of a real system, avoiding expensive or fragile components, emulator test rigs repre-
371
sent a possible solution. As demonstrated in previous papers [24,25], these facilities, designed to investigate hybrid system performance without expensive fuel cell stacks, are able to analyse fundamental issues [24] (e.g. component integration and control problems). While the plant developed at U.S. DOE–NETL is designed to emulate SOFC thermal behaviour with a combustor controlled by a fuel cell real-time model [24], the facility at the German Aerospace Centre (DLR) includes a water cooling system to emulate (with the combustor operation) the real stack exhaust gas composition [25]. They can operate at critical conditions without ceramic materials, carrying out tests with more flexibility and without expensive damages in case of failures while exploring very dangerous operative conditions, where existing control systems may be not very effective. A new SOFC hybrid system emulator has been developed at the University of Genoa with EU funding [27,28] including a complete (both cathode and anode sides) high temperature fuel cell system emulator [26]. This experimental facility has been used to analyse SOFC cathodic–anodic interactions from a fluid dynamic and thermal point of view. Attention has been focused on the differential pressure between the cell sides, on anodic recirculation performance (ejector recirculation factor and pressure increase) and on gas turbine performance behaviour. 2. Test rig The hybrid system emulator is based on a modified Turbec T100 microturbine (mGT) connected to a fuel cell emulator [26]. The rig is based on a modular vessel for cathodic side emulation [26] and an anodic recirculation system [28]. This device was developed to carry out tests on hybrid system configurations without the risks and costs related to the presence of an actual fuel cell stack. Even if no ceramic material is included in the rig, it is possible (as demonstrated in [24,25]) to perform significant tests on mGT/SOFC coupling (e.g. start-up and shutdown phases, cathode–anode fluid dynamic interaction) for phenomenon analysis and control system development purposes. This plant is the only hybrid system emulator rig equipped with an anodic recirculation currently in existence.
Fig. 3. Plant layout and main probes.
372
M.L. Ferrari, A.F. Massardo / Applied Energy 105 (2013) 369–379
Fig. 4. Image of the hybrid system emulator.
The whole test rig and the probe locations (used in this work) are shown in Fig. 3. The fuel cell emulator is coupled to the gas turbine through a set of connection pipes (Fig. 4). They were designed to manage the air flows (through VM, VR, VO valves of Fig. 3) and to measure the mass flow rate values [26]. The gas turbine is a Turbec T100 PH Series 3. Two different mGT control strategies have been used: (i) gas turbine operating connected to the electrical grid with a constant TOT control system [27], (ii) gas turbine in stand-alone mode with a constant rotational speed control system [27]. The details of its components, control system, and main design property values have been described in previous papers [26–29]. The T100 modifications for connection to external components have also been presented and discussed [26–28]. The fuel cell emulator was designed based on a hybrid system power of about 450 kW and a system efficiency of 59%. It is composed of two vessels (cathodic and anodic sides) sized to emulate a planar SOFC stack [13,14]. It was designed with a similitude approach (see [28] for details) based on the stack of Rolls-Royce Fuel Cell Systems [13,14]. The dimensions were based on a fuel cell size consistent with the gas turbine mass flow rate (0.8 kg/s at design conditions) [28]. The modular cathodic vessel is composed of two collector pipes, connected at gas turbine recuperator outlet [26] and combustor inlet respectively (Fig. 3), and four module pipes connected to both collectors (Fig. 4). Both collectors and module pipes have a nominal diameter of 350 mm and their total length is around 34 m. Since the modular characteristic of the cathodic vessel enables it to emulate fuel cells of different dimensions, the hybrid system emulation shown here is carried out with the maximum volume size (it
is about 3.2 m3) designed to emulate a 450 kW (electrical power) hybrid system [28]. The anodic recirculation loop is composed of a compressed air line (for fuel flow emulation), an anodic ejector, and an anodic vessel. The compressed air line is based on a 15 kW compressor to supply up to 20 g/s of air in the ejector primary duct. This air line is also equipped with an air dryer, and essential probes to measure the following properties: mass flow rate (MP with Fig. 3 nomenclature), pressure (PEjP1), and temperature (TEjP1). The anodic ejector design was based on previous validated works [15,23]. It is a typical device used in the anodic sides of hybrid systems to generate recirculation flow through SOFC ducts and to control the Steam-to-Carbon Ratio. Its volume is consistent with the SOFC cathodic size (about 0.8 m3). In addition, to obtain good thermal emulation on the anodic side, part of the anodic loop was inserted into the cathodic volume. This approach ensures that cathodic and anodic sides are thermally coupled, as in a real stack. For this reason, two innovative heat exchangers connected to the anodic loop were designed and installed inside the cathodic collectors. This plant detail is shown in Fig. 5. The test rig is managed by a suitable data acquisition and control system (see Fig. 6 for the main panel), developed by the Authors using LabVIEWTM software (see [26–28] for details). However, the original control system of the gas turbine was maintained to ensure that the turbine always is operated under safe conditions. A suitable interface was developed to obtain all the machine signals necessary for control and performance evaluation (e.g. turbine outlet temperature and rotational speed) in the LabVIEWTM software. Since room temperature is a significant variable for hybrid system operation [20], a specific further system was developed to manage the machine compressor inlet temperature (TC1 with Fig. 3 nomenclature).Three air/water heat exchangers [27] installed at the compressor intakes (‘‘Ex’’ components with Fig. 3 nomenclature) were used to control TC1 through cooling water. In addition, an absorption cooler was installed to generate chilled water exploiting machine co-generation thermal energy [30] obtained through the ‘‘WHEx’’ component. This allowed us to operate the rig at 15 °C (288.15 K) TC1 values even during the summer season. To complete the hybrid system emulation, a real-time model was developed in MatlabÒ–SimulinkÒ for components not physically present [30] in the rig (e.g. the stack). Through the Real-Time Windows Target tool and an UDP interface, this model is able to work in parallel with the plant. This model receives the values of air mass flow rate and temperature at the mGT combustor inlet level, and the gas turbine rotational speed from the plant. As stated in [30], the MatlabÒ–SimulinkÒ components are used to calculate TOT value. It is essential that the real gas turbine is controlled (with electrical load in stand-alone mode) to obtain the TOT value calculated by the model [30].
Fig. 5. Anodic loop details.
M.L. Ferrari, A.F. Massardo / Applied Energy 105 (2013) 369–379
373
Fig. 6. Main panel of control and acquisition system.
3. Cathode–anode interaction tests The cathode–anode interaction experimental results were obtained with the hybrid system emulator rig using both cathode and anode loops described in the previous paragraph. As stated above, given the significant influence of ambient temperature on system performance, the compressor inlet temperature control system was used to maintain TC1 value at 300 K. These interaction tests carried out on both cathode and anode sides were carried out for different operative gas turbine conditions, since the mGT commercial control system operates with two different strategies depending on the operating mode: Electrical grid-connected mode. The mGT operates with a variable speed control system: the turbine outlet temperature (TT2 using Fig. 3 nomenclature) is measured and kept constant via mGT combustor fuel variation. Stand-alone mode. The mGT operates at constant speed (mGT fuel is managed to maintain the rotational speed at 67,550 rpm). While in the electrical grid-connected mode the gas turbine generated power is simply delivered to the national electrical grid, in stand-alone mode a 100 kW resistor bank (managed by an inverter) is used to dispose of the generated power (emulation of a stand-alone grid).
3.1. Electrical grid-connected mode These tests were carried out for different net electrical load values starting from 20 kW (to avoid unstable condition risks, the gas turbine control system does not enable steady-state tests at less than 20 kW [28]) to the maximum load. At each electrical load value (20 kW, 40 kW, 60 kW, maximum load) different steady-state tests were carried out at various mass flow rate values in the anodic ejector primary duct (0 g/s, 5 g/s, 10 g/s, 15 g/s, 20 g/s) emulating fuel flow rate variation. In agreement with Fig. 3 nomenclature, the mass flow rate in the ejector primary duct is called MP. The case with MP = 0 g/s is not an operating condition significant for the operation of a real hybrid system, however it is reported here as a limit condition that can be obtained without anodic flow.
For the injection of this additional air mass flow rate the available maximum load increases with MP increase (it ranges from 71.9 kW at MP = 0 g/s to 74.1 kW at MP = 20 g/s). This behaviour is related to additional power available in the expander due to the additional mass flow rate and energy coming from the anodic circuit. Fig. 7 shows the gas turbine rotational speed and the anodic ejector recirculation ratio (the ratio of mass flow rate through the ejector secondary to the mass flow rate through the ejector primary ducts with the nomenclature shown in Fig. 2) related to its design value [28]. Anodic ejector performance is an essential aspect for hybrid systems equipped with these devices (high recirculation ratio values are essential to ensure thermal energy for reforming reaction and to avoid carbon deposition problems [15,23,31,32]). Starting from maximum load condition the rotational speed decreases with load generating a cathodic mass flow rate decrease to about 64% of its nominal value [28] at 20 kW load. Moreover, the ejector primary duct mass flow (related to hybrid system input energy) increase generates a slight rotational speed decrease (at constant load condition). This slight decrease is due to the additional mass flow rate available in the expander. Since this flow comes from the anodic circuit, and does not affect compressor consumed power, the power balance on the shaft is obtained at slightly lower rotational speed values (control system operates at constant TOT). In addition, Fig. 7 shows that the recirculation ratio increases with the load increase (at constant MP value). This is an effect of pressure increase due to gas turbine rotational speed increase. In details, as fully discussed in [15,23], a pressure increase at the secondary inlet duct generates, with the same ejector pressure rise, a recirculated mass flow rate increase. The ejector pressure rise (DPEj) shown in Fig. 8 is almost not significantly dependent by the gas turbine load. This is due to pressure and temperature trends at ejector secondary inlet (pressure variation effect is almost compensated by temperature trend). Since the air mass flow rate change generates a significant variation in heat exchange performance between cathode and anode sides, anodic duct temperatures increase with load increase. In details, even if the turbine outlet temperature is constant (TT2 = 918.15 K) during all the tests, the inlet temperature of the ejector secondary duct (TEjS1 in Fig. 3) increases from about 673 K at 20 kW load to about 703 K at maximum load. Moreover, the influence of MP on TEjS1 is almost negligible. Since the anodic
374
M.L. Ferrari, A.F. Massardo / Applied Energy 105 (2013) 369–379
Fig. 7. Anodic ejector recirculation ratio (F) referred to design value (F0) [28] and mGT rotational speed (ROT.SP.) in grid-connected mode. The ‘‘max kW’’ load ranges from 74.1 kW at MP = 20 g/s to 72.2 kW at MP = 5 g/s.
Fig. 8. Anodic ejector pressure rise (DPEj) in mGT grid-connected mode. The ‘‘max kW’’ load ranges from 74.1 kW at MP = 20 g/s to 72.2 kW at MP = 5 g/s.
side properties are not significant if MP is equal to 0 g/s, no anodic properties are shown in this condition in both Figs. 7 and 8. The effect of MP mass flow rate on the turbine side (stack emulator cathodic side) is shown in Figs. 9 and 10. The compressor discharge pressure (Fig. 9) measured downstream of the check valve (see Fig. 3 for the valve location) is reported considering the differ-
ent MP mass flow rate values. The trend of this property is mainly due to the rotational speed trend shown in Fig. 7. In detail, Fig. 9 shows a significant pressure increase with the load increase for the effect of rotational speed increase (also mass flow rate increase is significant). Another important aspect is present at constant load condition. This is a slight pressure decrease generated by the ejector primary duct mass flow rate (MP) increase. Since the machine operates with a constant TOT controller, the effect of rotational speed increase (shown in Fig. 7) is stronger than a pressurization effect due to additional air injection. For this reason, the effect of the MP mass flow rate increase on compressor outlet pressure is the slight decrease, shown in Fig. 9. Another effect of MP mass flow rate in ‘‘grid-connected’’ tests is shown in Fig. 10 for the recuperator outlet temperature at the cold side (TRC2) and the effectiveness (e). For the effectiveness peak trend, shown and discussed in [27], related to mass flow rate increase with load increase, both TRC2 and e (calculated with Eq. (1) and affected by a ±2% average accuracy value) have a peak trend with load increase (the maximum values are obtained with the tests at 60 kW instead with the ‘‘max kW’’ data). This peak trend is in accordance with previous theoretical works [33,34]. As already stated in [34], a significant impact of longitudinal conduction is affecting the effectiveness value when the recuperator is working at low flow condition. However, this effect is not as evident as it is in [34] because, to avoid unstable condition risks, the mGT control system does not enable steady-state tests at less than 20 kW. Moreover, at constant electrical load condition, Fig. 10 shows a TRC2 and e increase with MP increase. This effect is due to an increase in the convective heat transfer coefficient on the hot side of the recuperator, obtained with an additional air mass flow rate coming from the anodic side (TT2 value is constant in all the tests performed in grid-connected mode).
e¼
Fig. 9. Compressor outlet pressure (PRC1) in mGT grid-connected mode. The ‘‘max kW’’ load ranges from 74.1 kW at MP = 20 g/s to 71.9 kW at MP = 0 g/s.
TRC2 TRC1 TT2 TRC1
ð1Þ
3.1.1. Cathode–anode differential pressure As previously discussed, the cathode–anode differential pressure (DPCA) is an essential property that must be analyzed to prevent damages to the SOFC stack. As shown in Fig. 3 with the DPCA differential pressure sensor, this property is measured between the inlet pipes of cathodic and anodic vessels (inlet sections of the cathode and anode of the SOFC stack). This probe is used to measure differential pressure values between the cell sides. Fig. 11 shows that cathode–anode differential pressure value is null where the anodic side outlet flow is mixed with the cathodic flow (OGB
M.L. Ferrari, A.F. Massardo / Applied Energy 105 (2013) 369–379
375
Fig. 10. Recuperator outlet temperature (TRC2) and effectiveness (e) in mGT grid-connected mode. The ‘‘max kW’’ load ranges from 74.1 kW at MP = 20 g/s to 71.9 kW at MP = 0 g/s.
Fig. 11. Cathode–anode pressure layout.
Fig. 12. Cathode–anode differential pressure (DPCA) during the whole test in mGT grid-connected mode.
Fig. 13. Cathode–anode differential pressure (DPCA) in mGT grid-connected mode. The ‘‘max kW’’ load ranges from 74.1 kW at MP = 20 g/s to 71.9 kW at MP = 0 g/s.
component). However, upstream of the OGB (in the stack) the pressure values depend on the pressure losses in the cell sides and on the line performance: while (pcat)in is driven by compressor line performance, (pan)in is mainly a function of fuel mass flow rate and anodic ejector performance. For this reason, the differential pressure on the ceramic material of the stack may be signifi-
cant, especially at part-load conditions and during transient operations. The anodic recirculation loop (vessel and pipes) was designed and built to minimize the cathode–anode differential pressure (DPCA) at design conditions (gas turbine at maximum load and MP = 20 g/s) [28]. The design DPCA value is not null, but it is small (sustainable by the cell materials) as in a typical SOFC system. A DPCA design value under 10 mbar is an acceptable result (the DPCA design value in the test rig is about 7 mbar). Even if a null DPCA design value is desirable, some details in the emulator rig are not exactly evaluated in design calculations (e.g. gaskets, welding details, pipe roughness, bolts, and other installation details), as in a real hybrid plant, and assumptions used may represent real phenomena in a simplified approach. Fig. 12 shows the differential pressure values measured over time during a complete test carried out in electrical grid-connected mode. The initial part (before the ‘‘Test zone’’ of Fig. 12) was necessary to heat the rig pipes and to operate initial MP variations (the DPCA peaks obtained before time 22.40.00). The measurement values obtained after the heating period (inside the ‘‘Test zone’’) were then used to obtain the off-design DPCA trend shown in Fig. 13, required to manage hybrid system operations at part-load conditions. The negative peak at the end of the test is related to gas turbine shutdown phase: since the flow in the ejector primary duct (MP) was maintained at 20 g/s during shutdown, the pressure decrease in the cathodic side (due to the decrease in rotational speed) generates significant negative DPCA values (close to 30 mbar) to be considered for a real hybrid system design. The DPCA value increases with load (for cathodic mass flow rate increase) and with the decrease in MP value. This second trend has to be carefully managed when operating a hybrid system at part-load conditions, to avoid excessive mechanical stress on the cell sides. The maximum sustainable DPCA value depends on fuel cell technology and design. Even if this value for the RRFCS stack is not available for confidentiality reasons, the results obtained with the emulator rig are useful for reducing this stress if considered critical by manufacturer. For this reason, Fig. 13 is very interesting for control strategy development. At part-load condition (MP lower than its design value), the DPCA value can be reduced, decreasing machine load and, as a consequence, the rotational speed (this operation generates a decrease in the cathodic side mass flow rate value). Moreover, an air flow decrease is also necessary to control fuel cell temperature values (see [17] for details), avoiding excessive thermal stress on the materials. The control system must be tuned to manage air mass flow rate and avoid damage due to both mechanical and thermal stress. Tests carried out with emulators are essential to reduce the machine rotational speed at part-load conditions, maintaining both DPCA and fuel cell temperatures under safe conditions. An example of possible dangerous conditions caused by incorrect operations is shown in Fig. 14. A strong step increase in the
376
M.L. Ferrari, A.F. Massardo / Applied Energy 105 (2013) 369–379
Fig. 14. Critical variation of DPCA and compressor outlet pressure (PRC1) due to a surge event generated by MP fast increase (mGT grid-connected mode).
anodic ejector primary duct mass flow rate (from 0 g/s to about 16 g/s) produces a pressure increase in the anodic side (the DPCA value reaches 30.8 mbar after 4 s). In the meantime, this mass flow rate increase step produces an increase oscillation in compressor outlet pressure (PRC1) generating a surge event (detected by the machine vibration sensor with a 3:5 g m/s2 signal). The subsequent cathodic pressure decrease due to the surge cycle produces very dangerous DPCA peak shown in Fig. 14 (minimum value equal to 145.4 mbar), which in a real hybrid system with an SOFC stack would cause the fuel cell to break [35]. 3.2. Stand-alone mode These tests were carried out with different net electrical load values starting from gas turbine idle condition to 60 kW (this is the maximum load to operate tests with MP values up to 20 g/s without excessive stress on the gas turbine fuel compressor). As
for the grid-connected mode, at each electrical load value (idle, 20 kW, 40 kW, 60 kW) different steady-state tests were carried out at various mass flow rate values in the anodic ejector primary duct (MP = 0 g/s, 5 g/s, 10 g/s, 15 g/s, 20 g/s). The case with MP = 0 g/s is justified as a limit condition already discussed for the ‘‘grid-connected’’ tests. Fig. 15 shows the turbine outlet temperature (named TT2 with Fig. 3 nomenclature) and the anodic ejector recirculation ratio (related to its design value [28]) measured operating the gas turbine in stand-alone mode. The turbine outlet temperature significantly increases with load increase because rotational speed is maintained constant in this stand-alone mode. Moreover, the ejector primary duct mass flow increase generates a slight TT2 decrease (at constant load condition). This trend is due to the injection of a pre-compressed mass flow rate, just upstream of combustor inlet, that slightly increases expander power as discussed for the gridconnected mode. In addition, Fig. 15 shows that the recirculation ratio decreases with the load increase (at constant MP value). This is due to the ejector secondary inlet temperature (TEjS1) increase that is generated by the TT2 increase. The TEjS1 variation (from about 556 K at 0 kW load to about 932 K at 60 kW) is significantly higher than the pressure increase at the ejector secondary duct inlet. As shown in [15,28], a temperature increase at the secondary inlet duct of the anodic ejector generates a recirculated mass flow rate decrease (convergent nozzles at subsonic conditions [15]). Fig. 16 shows the ejector pressure rise (DPEj) measured during the tests in stand-alone mode. As already discussed for Fig. 8, this property (DPEj) is almost not significantly dependent by gas turbine load, because (for viscous losses) the pressure increase effect is compensated by temperature variation. Compressor outlet pressure is an important property that is essential to show cathode–anode side interaction related to MP flow injection. Since in this stand-alone mode machine rotational speed is constant, this increase in pressure (PRC1 in Fig. 17) with load increase is not as strong as the increase for grid-connected mode (Fig. 9). It is simply related to compressor pressure/flow map [26] trend at constant rotational speed. Moreover, a slight pressure increase (at constant load condition) with the mass flow rate increase in the ejector primary duct is evident in Fig. 17. In this case, the effect of additional air injection, at constant conditions, is a pressure increase. Since turbine rotational speed is constant (in stand-alone mode), the air injection at combustor inlet generates a slight additional pressure at compressor discharge level. The consequence of this trend is a dangerous decrease in surge margin. Although this aspect is not significant for the turbine in the tests
Fig. 15. Anodic ejector recirculation ratio (F) referred to design value (F0) [28] and mGT turbine outlet temperature (TT2) in stand-alone mode.
M.L. Ferrari, A.F. Massardo / Applied Energy 105 (2013) 369–379
377
Fig. 16. Anodic ejector pressure rise (DPEj) in mGT stand-alone mode. Fig. 19. Cathode–anode differential pressure (DPCA) in mGT stand-alone mode.
there is a decrease in effectiveness for the temperature increase (for electrical load increase) and an opposite behaviour for the test at 60 kW load (as shown in [27]) due to heat transfer coefficient performance. This increase obtained at 60 kW load it is not affecting the general decrease trend with load increase, because the data reported in [27] show that at 72.1 kW (for MP = 0 g/s) the effectiveness is lower than the value obtained for 40 kW load. For these tests, effectiveness is subject to an accuracy ranging from ±6% at 0 kW load to ±3% at 60 kW (see [27] for details). Moreover, at constant electrical load condition, TRC2 shows a slight decrease for TT2 decrease with MP increase. However, this TRC2 decrease is less significant than the TT2 trend and there is a strong e increase effect as MP increases. This is due to an increase in convective heat exchange coefficient on the recuperator hot side obtained with an additional air mass flow rate from the anodic side (as already discussed for the electrical grid-connected mode).
Fig. 17. Compressor outlet pressure (PRC1) in mGT stand-alone mode.
reported here, compressor outlet pressure increase must always be carefully managed for the safe operations of a hybrid system, especially if the turbine has a low surge margin. The recuperator outlet temperature at the cold side (TRC2) and the effectiveness (e) are reported in Fig. 18 for the stand-alone mode tests. TRC2 shows a typical increase trend with load due to TT2 increase. For the effectiveness trend obtained with the increase in the electrical load, the results shown in Fig. 18 are in agreement with the values discussed in [27]. In details, it can be observed that
3.2.1. Cathode–anode differential pressure Also at stand-alone conditions, the cathode–anode differential pressure (Fig. 19) is analyzed to prevent excessive stress on SOFC ceramic material. In this operating mode, the load influence on DPCA is almost negligible because the cathodic mass flow rate variation is not significant, as it is for the grid-connected mode (the rotational speed is constant during all the tests). Moreover, Fig. 19 shows a DPCA increase with the decrease in MP (for each load value), due to the decrease in mass flow rate in the anodic side. As discussed for the grid-connected mode, this aspect has to be considered for the development of control strategies for hybrid systems. The experimental tests reported here show that if used in a hybrid system, a constant rotational speed gas turbine generates further part-load constraints. In details, if rotational speed is constant the cathode–anode differential pressure has to be managed
Fig. 18. Recuperator outlet temperature (TRC2) and effectiveness (e) in mGT stand-alone mode.
378
M.L. Ferrari, A.F. Massardo / Applied Energy 105 (2013) 369–379
through alternative approaches (not always efficient from the turbine point of view). However, if this kind of turbine control strategy is mandatory (e.g., for safety constraints due to stand-alone operations [26]), bleed approaches or mass flow rate bypass solutions (see [24] for details) have to be designed and carefully managed at part-load conditions (to discharge or bypass the excessive amount of compressor outlet flow). The results discussed here show that, if the machine operates with constant speed, specific bleed or bypass strategies are necessary not only to control SOFC temperatures, but also to avoid excessive stress related to high DPCA values.
4. Conclusions This paper presents an experimental analysis of cathode–anode interaction for an SOFC based hybrid system. To avoid the risks and costs related to an actual fuel cell stack, the emulator rig developed at the University of Genoa was used for a broad series of experiments. As demonstrated in previous works [24,25] this kind of test rig (based on the coupling of a microturbine with a complete cathode–anode fuel cell emulator) is able to produce effects similar to real plant results. The following main results were obtained: The recirculation ratio (F) for the anodic ejector showed different trends. While for mGT grid-connected mode F increases when the load increases depending on MP (from about 2% at 5 g/s to about 25% at 20 g/s), in stand-alone mode F decreases with the load increase (almost constant decrement of about 10%). The rise in ejector pressure (DPEj) is not significantly dependent on mGT load for both grid-connected and stand-alone modes. At constant mGT electrical load, for grid-connected mode the increase in anodic ejector mass flow rate in the primary duct (MP) generates a decrease in rotational speed (average variation: 1.4%) while for stand-alone condition it is related to a TOT decrease (variation: 2.6%). At mGT stand-alone mode the air injection on the anodic side (MP) generates an increase in compressor outlet pressure (average variation: <1%) at constant electrical load. This trend is detrimental in terms of the surge margin. Even if it is not significant for these tests, this has to be managed during hybrid system operations (especially with turbines affected by low surge margin performance). From the thermal point of view the trend in recuperator effectiveness showed a slight increase (between 1.1% and 1.7%) for both grid-connected and stand-alone modes (because of the increased heat exchange coefficient due to the additional mass flow rate). Although the performance increase is small, this might be interesting for hybrid system operations. While for mGT grid-connected mode the load increase generates a significant increase in cathode–anode differential pressure (average variation: 12.4 mbar), for mGT stand-alone configuration the electrical load does not appear to significantly influence this property. For both grid-connected and stand-alone modes, cathode– anode differential pressure significantly increases (about 20 mbar) with the decrease in ejector primary duct mass flow. This aspect has to be carefully taken into account to avoid excessive mechanical stress on the fuel cell materials at partload conditions. These experiments are very useful to develop a suitable control strategy to couple the decrease in fuel mass flow rate (emulated with MP) with the decrease in cathode side air.
As discussed in the text this emulator is also equipped with a real-time model [30] for components that are not physically present in the laboratory (e.g. the fuel cell stack). The approach based on hardware–software coupling (see [30] for details) will be essential to develop further tests on cathode–anode interaction during actual operations, focusing attention on critical phases (e.g. startup, shutdown and load changes).
Acknowledgements This work was funded by Felicitas (TIP4-CT-2005-516270) and LARGE-SOFC (No. 019739) European Projects. Additionally, the authors would like to thank Dr. Matteo Pascenti of TPG for his support in the experimental activity. The results presented and discussed here are on the sole responsibility of the authors and do not represent the views of the company cited above.
References [1] International Energy Agency. Key World Energy Statistics, 2006. [2] Kathib H, Energy considerations global warming perspectives, World Energy Council. Energy Permanent Monitoring Panel of the World Federation of Sciences, 2004. [3] Goodstein ES. Economics and the environment. Englewood Cliffs (NJ): Prentice-Hall; 1999. [4] Quirion P. Complying with the kyoto protocol under uncertainty: Taxes or tradable permits? Energy Policy 2010;38(9):5166–73. [5] Veyo S, Shockling LA, Dederer JT, Gillett JE, Lundberg WL. Tubular solid oxide fuel cell/gas turbine hybrid cycle power systems: status. J Eng Gas Turbines Power 2002;124:845–9. [6] Birgersson KE, Balaya P, Chou SK, Yan J. Energy solutions for a sustainable world. Appl Energy 2012;90:1–2. [7] Andersson M, Yuan J, Sundén B. Review on modeling development for multiscale chemical reactions coupled transport phenomena in solid oxide fuel cells. Appl Energy 2010;87:1461–76. [8] Richards GA, McMillian MM, Gemmen RS, Rogers WA, Cully SR. Issues for lowemission, fuel-flexible power systems. Progr Energy Combust Sci 2001;27:141–69. [9] Nanaeda K, Mueller F, Brouwer J, Samuelsen S. Dynamic modeling and evaluation of solid oxide fuel cell – combined heat and power system operating strategies. J Power Sources 2010;195:3176–85. [10] Riensche E, Stimming U, Unverzagt G. Optimization of a 200 kW SOFC cogeneration power plant: Part I: Variation of process parameters. J Power Sources 1998;72:251–6. [11] Mueller F, Brouwer J, Jabbari F, Samuelsen G. Dynamic simulation of an integrated solid oxide fuel cell system including current-based fuel flow control. J Fuel Cell Sci Technol 2006;3:144–55. [12] Calise F, Ferruzzi G, Vanoli L. Parametric exergy analysis of a tubular solid oxide fuel cell (SOFC) stack through finite-volume model. Appl Energy 2009;86:2401–10. [13] Agnew GD, Bozzolo M, Moritz RR, Berenyi S. The design and integration of the Rolls-Royce Fuel Cell Systems 1MW SOFC. ASME Paper GT2005-69122, 2005. [14] Trasino F, Bozzolo M, Magistri L, Massardo AF. Modeling and performance analysis of the Rolls-Royce Fuel Cell systems limited: 1 MW plant. J Eng Gas Turbines Power 2011;133. p. 021701_1–11. [15] Ferrari ML, Pascenti M, Massardo AF. Ejector model for high temperature fuel cell hybrid systems: experimental validation at steady-state and dynamic conditions. J Fuel Cell Sci Technol 2008;5. p. 041005_1–7. [16] Ghigliazza F, Traverso A, Massardo AF, Wingate J, Ferrari ML. Generic real-time modeling of solid oxide fuel cell hybrid systems. J Fuel Cell Sci Technol 2009;6. p. 021312_1–7. [17] Ferrari ML. Solid oxide fuel cell hybrid system: control strategy for stand-alone configurations. J Power Sources, Elsevier 2011;196(5):2682–90. [18] Shibata S, Urata T, Sugawara Y, Umeda T, Morita J, Hatoh K, Kitano Y. Fuel cell system with differential pressure control. U.S. Patent US7691510B2, 2010. [19] Limbeck U, Mazzotta CS, Pasera U, Schmalzriedt S, Wiesner P. Method and apparatus for controlling the differential pressure in a fuel cell. U.S. Patent US2007/0184319A1, 2007. [20] McLarty D, Kuniba Y, Brouwer J, Samuelsen S. Experimental and theoretical evidence for control requirements in solid oxide fuel cell gas turbine hybrid systems. J Power Sources 2012;209:195–203. [21] Massardo AF, Lubelli F. Internal reforming solid oxide fuel cell-gas turbine combined cycles (IRSOFC-GT) – Part I: Cell model and cycle thermodynamic analysis. J Eng Gas Turbines Power 2000;122:27–35. [22] Massardo AF, Magistri L. Internal reforming solid oxide fuel cell gas turbine combined cycles (IRSOFC-GT) – Part II: Exergy and thermoeconomic analyses. J Eng Gas Turbines Power 2003;125:67–74.
M.L. Ferrari, A.F. Massardo / Applied Energy 105 (2013) 369–379 [23] Ferrari ML, Traverso A, Magistri L, Massardo AF. Influence of the anodic recirculation transient behaviour on the SOFC hybrid system performance. J Power Sources 2005;149:22–32. [24] Shelton M, Celik I, Liese E, Tucker D. A study in the process modeling of the startup of fuel cell/gas turbine hybrid systems. J Eng Gas Turbines Power 2010;132. p. 012301_1–8. [25] Hohloch M, Widenhorn A, Lebküchner D, Panne T, Aigner M. Micro gas turbine test rig for hybrid power plant application. ASME Paper GT2008-50443, 2008. [26] Ferrari ML, Pascenti M, Magistri L, Massardo AF. Hybrid system test rig: startup and shutdown physical emulation. J Fuel Cell Sci Technol 2010;7. p. 021005_1–7. [27] Ferrari ML, Pascenti M, Magistri L, Massardo AF. Micro gas turbine recuperator: steady-state and transient experimental investigation. J Eng for Gas Turbines Power 2010;132. p. 022301_1–8. [28] Ferrari ML, Pascenti M, Magistri L, Massardo AF. MGT/HTFC hybrid system emulator test rig: experimental investigation on the anodic recirculation system. J Fuel Cell Sci Technol 2011;8. p. 021012_1–9.
379
[29] Ferrari ML, Pascenti M, Traverso AN, Massardo AF. Hybrid system test rig: chemical composition emulation with steam injection. Appl Energy 2012;97:809–15. [30] Ferrari ML, Pascenti M. Flexible micro gas turbine rig for tests on advanced energy systems, Advance in gas turbine technology. Rijeka, Croatia: INTECH; 2011. ISBN 978-953-308-68-4, p. 89–114, [chapter 5]. [31] Marsano F, Magistri L, Massardo AF. Ejector performance influence on a solid oxide fuel cell anodic recirculation system. J Power Sources 2004;129:216–28. [32] Ferrari ML, Bernardi D, Massardo AF. Design and testing of ejectors for high temperature fuel cell hybrid systems. J Fuel Cell Sci Technol 2006;3:284–91. [33] McDonald CF. Heat recovery exchanger technology for very small gas turbines. Int J Turbo Jet Engines 1996;13:239–61. [34] McDonald CF. Recuperator considerations for future high efficiency microturbines. J Appl Therm Energy 2003;23:1453–87. [35] Ferrari ML, Pascenti M, Bertone R, Magistri L. Hybrid simulation facility based on commercial 100 kWe micro gas turbine. J Fuel Cell Sci Technol 2009;6. p. 031008_1–8.
Contents lists available at SciVerse ScienceDirect
Applied Energy journal homepage: www.elsevier.com/locate/apenergy
Cathode–anode side interaction in SOFC hybrid systems Mario L. Ferrari ⇑, Aristide F. Massardo Thermochemical Power Group (TPG), DIME, University of Genoa, Italy
h i g h l i g h t s " Cathode–anode interaction issues. " Hybrid system emulator with anodic recirculation. " Grid-connected and stand-alone tests. " Cathode–anode differential pressure. " Control system aspects.
a r t i c l e
i n f o
Article history: Received 31 July 2012 Received in revised form 3 January 2013 Accepted 8 January 2013 Available online 9 February 2013 Keywords: SOFC hybrid system Microturbine Test rig Cathode–anode interaction
a b s t r a c t Cathode–anode interaction, mainly based on cathode versus anode volume influence, recirculation performance, and turbomachinery integration, is an important issue for pressurised SOFC hybrid systems, and this aspect must be carefully considered to prevent fuel cell ceramic material failures through a reliable control system. Over the last 10 years, several theoretical analyses of this issue have been carried out at the University of Genoa. These interaction studies have been analysed and an experimental approach (for model validation, system development and prototype design activities) has been applied using emulator facilities or real plants. In particular, general hybrid system layouts based on the coupling of pressurized SOFC stacks of different geometries (planar, tubular, etc.) with a gas turbine bottoming cycle have been investigated using the hybrid system emulator facility of the University of Genoa. The experimental results are focused on the interaction between gas turbine and anodic circuit and on cathode–anode differential pressure behaviour for design, off-design and transient hybrid system operative conditions. The information obtained in these tests is essential to understand the main features of the variables that drive the phenomena and to design a suitable control system that can mitigate the differential pressure values during all plant operating conditions. Ó 2013 Elsevier Ltd. All rights reserved.
1. Introduction The increase in demand for electrical energy [1], coupled with a heightened awareness of environmental pollution problems [2] related to power system emissions, provides a strong incentive for researchers to develop more efficient plants for power generation. In addition, since conventional fossil fuel generation activity is the major source of greenhouse gas emissions [3], this field of industry certainly needs to improve to meet international targets [4]. In this scenario, high temperature fuel cell hybrid systems are a possible effective solution to environmental and efficiency issues [5–7]. This technology is based on the coupling of a fuel cell stack (able to convert the chemical energy of fuel directly into electricity without the efficiency limitations typical of thermodynamic cycles) ⇑ Corresponding author. E-mail addresses: [email protected] (M.L. Ferrari), [email protected] (A.F. Massardo). 0306-2619/$ - see front matter Ó 2013 Elsevier Ltd. All rights reserved. http://dx.doi.org/10.1016/j.apenergy.2013.01.029
with a conventional gas turbine. In addition to efficiency benefits, hybrid systems produce low pollutant [8] exhaust gases at temperature values useful for co-generative applications [8–10]. Two pressurised hybrid system general layouts (Fig. 1), based on solid oxide fuel cell (SOFC) technology developed by Siemens Westinghouse [5,11,12] and Rolls-Royce Fuel Cell Systems [13,14], are considered in this paper. The fuel cell, usually located between compressor and turbine (pressurised solution), is equipped with an ejector based anodic recirculation and an offgas burner. Fuel is fed into the primary duct of an anodic ejector [15] (see Fig. 2 for ejector nomenclature details). A reforming section may be present upstream of the anode inlet [11] or located internally in the fuel cell [14]. In the ‘‘Layout A’’ a standard recuperator is included to pre-heat the air flow upstream of the fuel cell inlet exploiting turbine exhaust thermal energy [15]. This configuration is an interesting and efficient solution for an SOFC stack equipped with an air pre-heating heat exchanger (or tube), such as the tubular fuel cells developed by Siemens Westinghouse
370
M.L. Ferrari, A.F. Massardo / Applied Energy 105 (2013) 369–379
Nomenclature F N TOT V
Acronyms AC alternate current DC direct current Ex heat exchanger mGT micro-gas turbine OGB off-gas burner REC recuperator SOFC solid oxide fuel cell TPG thermochemical power group UDP user datagram protocol WHEx water heat exchanger (for co-generation)
recirculation ratio () rotational speed (rpm) turbine outlet temperature (K) volume (m3)
Greek symbols e effectiveness Subscripts 0 on-design an anodic cat cathodic in inlet out outlet
Variables g 9.81 (m/s2)
[5,11]. ‘‘Layout B’’ is based on a cathodic recirculation loop carried out through a cathodic ejector [15]. This recirculation layout is less efficient for air heating but useful to avoid recuperator costs and high temperature heat exchanger constraints (and also very interesting for the thermal control of the stack [14,16]). Cathode and anode sides volume sizes are different for the two schemes shown in Fig. 1. This is related to the different fuel cell technology: while the tubular cell of the ‘‘Layout A’’ has an anodic volume significantly higher than the cathodic one (cathode/anode volume ratio around 0.5 [17]), the reverse (cathode/anode volume ratio around 4 [16]) applies to fuel cells not equipped with high temperature air pre-heating systems (e.g. planar SOFCs). Moreover, the two layouts have different values for pressure loss in the cell sides. For this reason, even if cathode and anode flows are mixed in the OGB in both layouts, different pressure values are possible in the upstream ducts between cathode and anode sides, especially during hybrid system part-load and time-dependent operations. Thermo-chemical aspects in time-dependent modes (e.g. startup, shutdown and load change phases) and control performance
depend on dynamic constants related to volume dimension (besides being influenced by ejector and machine performance, which are essential for the main influence properties: ejector recirculation ratio, ejector pressure increase, compressor delivery pressure). Cathode–anode interaction analysis (mainly the results of the pressure difference between fuel cell sides) is important to understand the effects on the bottoming cycle gas turbine components and for SOFC material safe management. In details, anodic side performance affects the cathode side due to the additional mass flow rate from fuel injection and the related modification of energy input. This generates possibly significant variations in cathodic side properties, such as machine rotational speed (or turbine outlet temperature), compressor outlet pressure and recuperator performance (see ‘‘Layout A’’ in Fig. 1) or cathodic recirculation aspects (‘‘Layout B’’ in Fig. 1). Since this behaviour can produce dangerous effects (e.g. surge events for compressor outlet pressure increase) at off-design conditions (or during transient operations), a complete analysis is essential.
Fig. 1. Hybrid systems: general layouts.
M.L. Ferrari, A.F. Massardo / Applied Energy 105 (2013) 369–379
Fig. 2. Ejector details.
Cell cathode–anode differential pressure behaviour is a very important aspect of this interaction. Apart from some patents [18,19] based on cathode–anode differential pressure management for low temperature fuel cell systems (turbomachinery effects are quite limited), no detailed theoretical investigations and experimental data have been published on this topic for SOFC hybrid systems. Moreover, no experimental data have been published on cathode–anode differential pressure trends at part-load and time-dependent conditions. On the other hand, many authors have been involved in theoretical activities analysing whole hybrid systems at different operative conditions [14,17,20–22], including control system issues [17]. For instance, cathode–anode interaction aspects were analysed in a previous model-based activity [23], in which the Authors calculated a 15–20 mbar cathode–anode pressure increase for each 10% power decrease (from nominal load value) if no control system is included for air flow. A subsequent publication [17] demonstrated how this mechanical stress could be avoided by controlling the gas turbine rotational speed. Even if the cathode–anode stress problem can be studied at stack level to increase ceramic material resistance performance, system level studies are needed to analyse the stress trend during part-load and time-dependent conditions. Since this kind of data is not available in open literature (manufacturers do not usually provide details on these aspects), a broad experimental support is mandatory to completely understand the phenomenon, validate models, and undertake prototype development activities. To generate experimental data representative of a real system, avoiding expensive or fragile components, emulator test rigs repre-
371
sent a possible solution. As demonstrated in previous papers [24,25], these facilities, designed to investigate hybrid system performance without expensive fuel cell stacks, are able to analyse fundamental issues [24] (e.g. component integration and control problems). While the plant developed at U.S. DOE–NETL is designed to emulate SOFC thermal behaviour with a combustor controlled by a fuel cell real-time model [24], the facility at the German Aerospace Centre (DLR) includes a water cooling system to emulate (with the combustor operation) the real stack exhaust gas composition [25]. They can operate at critical conditions without ceramic materials, carrying out tests with more flexibility and without expensive damages in case of failures while exploring very dangerous operative conditions, where existing control systems may be not very effective. A new SOFC hybrid system emulator has been developed at the University of Genoa with EU funding [27,28] including a complete (both cathode and anode sides) high temperature fuel cell system emulator [26]. This experimental facility has been used to analyse SOFC cathodic–anodic interactions from a fluid dynamic and thermal point of view. Attention has been focused on the differential pressure between the cell sides, on anodic recirculation performance (ejector recirculation factor and pressure increase) and on gas turbine performance behaviour. 2. Test rig The hybrid system emulator is based on a modified Turbec T100 microturbine (mGT) connected to a fuel cell emulator [26]. The rig is based on a modular vessel for cathodic side emulation [26] and an anodic recirculation system [28]. This device was developed to carry out tests on hybrid system configurations without the risks and costs related to the presence of an actual fuel cell stack. Even if no ceramic material is included in the rig, it is possible (as demonstrated in [24,25]) to perform significant tests on mGT/SOFC coupling (e.g. start-up and shutdown phases, cathode–anode fluid dynamic interaction) for phenomenon analysis and control system development purposes. This plant is the only hybrid system emulator rig equipped with an anodic recirculation currently in existence.
Fig. 3. Plant layout and main probes.
372
M.L. Ferrari, A.F. Massardo / Applied Energy 105 (2013) 369–379
Fig. 4. Image of the hybrid system emulator.
The whole test rig and the probe locations (used in this work) are shown in Fig. 3. The fuel cell emulator is coupled to the gas turbine through a set of connection pipes (Fig. 4). They were designed to manage the air flows (through VM, VR, VO valves of Fig. 3) and to measure the mass flow rate values [26]. The gas turbine is a Turbec T100 PH Series 3. Two different mGT control strategies have been used: (i) gas turbine operating connected to the electrical grid with a constant TOT control system [27], (ii) gas turbine in stand-alone mode with a constant rotational speed control system [27]. The details of its components, control system, and main design property values have been described in previous papers [26–29]. The T100 modifications for connection to external components have also been presented and discussed [26–28]. The fuel cell emulator was designed based on a hybrid system power of about 450 kW and a system efficiency of 59%. It is composed of two vessels (cathodic and anodic sides) sized to emulate a planar SOFC stack [13,14]. It was designed with a similitude approach (see [28] for details) based on the stack of Rolls-Royce Fuel Cell Systems [13,14]. The dimensions were based on a fuel cell size consistent with the gas turbine mass flow rate (0.8 kg/s at design conditions) [28]. The modular cathodic vessel is composed of two collector pipes, connected at gas turbine recuperator outlet [26] and combustor inlet respectively (Fig. 3), and four module pipes connected to both collectors (Fig. 4). Both collectors and module pipes have a nominal diameter of 350 mm and their total length is around 34 m. Since the modular characteristic of the cathodic vessel enables it to emulate fuel cells of different dimensions, the hybrid system emulation shown here is carried out with the maximum volume size (it
is about 3.2 m3) designed to emulate a 450 kW (electrical power) hybrid system [28]. The anodic recirculation loop is composed of a compressed air line (for fuel flow emulation), an anodic ejector, and an anodic vessel. The compressed air line is based on a 15 kW compressor to supply up to 20 g/s of air in the ejector primary duct. This air line is also equipped with an air dryer, and essential probes to measure the following properties: mass flow rate (MP with Fig. 3 nomenclature), pressure (PEjP1), and temperature (TEjP1). The anodic ejector design was based on previous validated works [15,23]. It is a typical device used in the anodic sides of hybrid systems to generate recirculation flow through SOFC ducts and to control the Steam-to-Carbon Ratio. Its volume is consistent with the SOFC cathodic size (about 0.8 m3). In addition, to obtain good thermal emulation on the anodic side, part of the anodic loop was inserted into the cathodic volume. This approach ensures that cathodic and anodic sides are thermally coupled, as in a real stack. For this reason, two innovative heat exchangers connected to the anodic loop were designed and installed inside the cathodic collectors. This plant detail is shown in Fig. 5. The test rig is managed by a suitable data acquisition and control system (see Fig. 6 for the main panel), developed by the Authors using LabVIEWTM software (see [26–28] for details). However, the original control system of the gas turbine was maintained to ensure that the turbine always is operated under safe conditions. A suitable interface was developed to obtain all the machine signals necessary for control and performance evaluation (e.g. turbine outlet temperature and rotational speed) in the LabVIEWTM software. Since room temperature is a significant variable for hybrid system operation [20], a specific further system was developed to manage the machine compressor inlet temperature (TC1 with Fig. 3 nomenclature).Three air/water heat exchangers [27] installed at the compressor intakes (‘‘Ex’’ components with Fig. 3 nomenclature) were used to control TC1 through cooling water. In addition, an absorption cooler was installed to generate chilled water exploiting machine co-generation thermal energy [30] obtained through the ‘‘WHEx’’ component. This allowed us to operate the rig at 15 °C (288.15 K) TC1 values even during the summer season. To complete the hybrid system emulation, a real-time model was developed in MatlabÒ–SimulinkÒ for components not physically present [30] in the rig (e.g. the stack). Through the Real-Time Windows Target tool and an UDP interface, this model is able to work in parallel with the plant. This model receives the values of air mass flow rate and temperature at the mGT combustor inlet level, and the gas turbine rotational speed from the plant. As stated in [30], the MatlabÒ–SimulinkÒ components are used to calculate TOT value. It is essential that the real gas turbine is controlled (with electrical load in stand-alone mode) to obtain the TOT value calculated by the model [30].
Fig. 5. Anodic loop details.
M.L. Ferrari, A.F. Massardo / Applied Energy 105 (2013) 369–379
373
Fig. 6. Main panel of control and acquisition system.
3. Cathode–anode interaction tests The cathode–anode interaction experimental results were obtained with the hybrid system emulator rig using both cathode and anode loops described in the previous paragraph. As stated above, given the significant influence of ambient temperature on system performance, the compressor inlet temperature control system was used to maintain TC1 value at 300 K. These interaction tests carried out on both cathode and anode sides were carried out for different operative gas turbine conditions, since the mGT commercial control system operates with two different strategies depending on the operating mode: Electrical grid-connected mode. The mGT operates with a variable speed control system: the turbine outlet temperature (TT2 using Fig. 3 nomenclature) is measured and kept constant via mGT combustor fuel variation. Stand-alone mode. The mGT operates at constant speed (mGT fuel is managed to maintain the rotational speed at 67,550 rpm). While in the electrical grid-connected mode the gas turbine generated power is simply delivered to the national electrical grid, in stand-alone mode a 100 kW resistor bank (managed by an inverter) is used to dispose of the generated power (emulation of a stand-alone grid).
3.1. Electrical grid-connected mode These tests were carried out for different net electrical load values starting from 20 kW (to avoid unstable condition risks, the gas turbine control system does not enable steady-state tests at less than 20 kW [28]) to the maximum load. At each electrical load value (20 kW, 40 kW, 60 kW, maximum load) different steady-state tests were carried out at various mass flow rate values in the anodic ejector primary duct (0 g/s, 5 g/s, 10 g/s, 15 g/s, 20 g/s) emulating fuel flow rate variation. In agreement with Fig. 3 nomenclature, the mass flow rate in the ejector primary duct is called MP. The case with MP = 0 g/s is not an operating condition significant for the operation of a real hybrid system, however it is reported here as a limit condition that can be obtained without anodic flow.
For the injection of this additional air mass flow rate the available maximum load increases with MP increase (it ranges from 71.9 kW at MP = 0 g/s to 74.1 kW at MP = 20 g/s). This behaviour is related to additional power available in the expander due to the additional mass flow rate and energy coming from the anodic circuit. Fig. 7 shows the gas turbine rotational speed and the anodic ejector recirculation ratio (the ratio of mass flow rate through the ejector secondary to the mass flow rate through the ejector primary ducts with the nomenclature shown in Fig. 2) related to its design value [28]. Anodic ejector performance is an essential aspect for hybrid systems equipped with these devices (high recirculation ratio values are essential to ensure thermal energy for reforming reaction and to avoid carbon deposition problems [15,23,31,32]). Starting from maximum load condition the rotational speed decreases with load generating a cathodic mass flow rate decrease to about 64% of its nominal value [28] at 20 kW load. Moreover, the ejector primary duct mass flow (related to hybrid system input energy) increase generates a slight rotational speed decrease (at constant load condition). This slight decrease is due to the additional mass flow rate available in the expander. Since this flow comes from the anodic circuit, and does not affect compressor consumed power, the power balance on the shaft is obtained at slightly lower rotational speed values (control system operates at constant TOT). In addition, Fig. 7 shows that the recirculation ratio increases with the load increase (at constant MP value). This is an effect of pressure increase due to gas turbine rotational speed increase. In details, as fully discussed in [15,23], a pressure increase at the secondary inlet duct generates, with the same ejector pressure rise, a recirculated mass flow rate increase. The ejector pressure rise (DPEj) shown in Fig. 8 is almost not significantly dependent by the gas turbine load. This is due to pressure and temperature trends at ejector secondary inlet (pressure variation effect is almost compensated by temperature trend). Since the air mass flow rate change generates a significant variation in heat exchange performance between cathode and anode sides, anodic duct temperatures increase with load increase. In details, even if the turbine outlet temperature is constant (TT2 = 918.15 K) during all the tests, the inlet temperature of the ejector secondary duct (TEjS1 in Fig. 3) increases from about 673 K at 20 kW load to about 703 K at maximum load. Moreover, the influence of MP on TEjS1 is almost negligible. Since the anodic
374
M.L. Ferrari, A.F. Massardo / Applied Energy 105 (2013) 369–379
Fig. 7. Anodic ejector recirculation ratio (F) referred to design value (F0) [28] and mGT rotational speed (ROT.SP.) in grid-connected mode. The ‘‘max kW’’ load ranges from 74.1 kW at MP = 20 g/s to 72.2 kW at MP = 5 g/s.
Fig. 8. Anodic ejector pressure rise (DPEj) in mGT grid-connected mode. The ‘‘max kW’’ load ranges from 74.1 kW at MP = 20 g/s to 72.2 kW at MP = 5 g/s.
side properties are not significant if MP is equal to 0 g/s, no anodic properties are shown in this condition in both Figs. 7 and 8. The effect of MP mass flow rate on the turbine side (stack emulator cathodic side) is shown in Figs. 9 and 10. The compressor discharge pressure (Fig. 9) measured downstream of the check valve (see Fig. 3 for the valve location) is reported considering the differ-
ent MP mass flow rate values. The trend of this property is mainly due to the rotational speed trend shown in Fig. 7. In detail, Fig. 9 shows a significant pressure increase with the load increase for the effect of rotational speed increase (also mass flow rate increase is significant). Another important aspect is present at constant load condition. This is a slight pressure decrease generated by the ejector primary duct mass flow rate (MP) increase. Since the machine operates with a constant TOT controller, the effect of rotational speed increase (shown in Fig. 7) is stronger than a pressurization effect due to additional air injection. For this reason, the effect of the MP mass flow rate increase on compressor outlet pressure is the slight decrease, shown in Fig. 9. Another effect of MP mass flow rate in ‘‘grid-connected’’ tests is shown in Fig. 10 for the recuperator outlet temperature at the cold side (TRC2) and the effectiveness (e). For the effectiveness peak trend, shown and discussed in [27], related to mass flow rate increase with load increase, both TRC2 and e (calculated with Eq. (1) and affected by a ±2% average accuracy value) have a peak trend with load increase (the maximum values are obtained with the tests at 60 kW instead with the ‘‘max kW’’ data). This peak trend is in accordance with previous theoretical works [33,34]. As already stated in [34], a significant impact of longitudinal conduction is affecting the effectiveness value when the recuperator is working at low flow condition. However, this effect is not as evident as it is in [34] because, to avoid unstable condition risks, the mGT control system does not enable steady-state tests at less than 20 kW. Moreover, at constant electrical load condition, Fig. 10 shows a TRC2 and e increase with MP increase. This effect is due to an increase in the convective heat transfer coefficient on the hot side of the recuperator, obtained with an additional air mass flow rate coming from the anodic side (TT2 value is constant in all the tests performed in grid-connected mode).
e¼
Fig. 9. Compressor outlet pressure (PRC1) in mGT grid-connected mode. The ‘‘max kW’’ load ranges from 74.1 kW at MP = 20 g/s to 71.9 kW at MP = 0 g/s.
TRC2 TRC1 TT2 TRC1
ð1Þ
3.1.1. Cathode–anode differential pressure As previously discussed, the cathode–anode differential pressure (DPCA) is an essential property that must be analyzed to prevent damages to the SOFC stack. As shown in Fig. 3 with the DPCA differential pressure sensor, this property is measured between the inlet pipes of cathodic and anodic vessels (inlet sections of the cathode and anode of the SOFC stack). This probe is used to measure differential pressure values between the cell sides. Fig. 11 shows that cathode–anode differential pressure value is null where the anodic side outlet flow is mixed with the cathodic flow (OGB
M.L. Ferrari, A.F. Massardo / Applied Energy 105 (2013) 369–379
375
Fig. 10. Recuperator outlet temperature (TRC2) and effectiveness (e) in mGT grid-connected mode. The ‘‘max kW’’ load ranges from 74.1 kW at MP = 20 g/s to 71.9 kW at MP = 0 g/s.
Fig. 11. Cathode–anode pressure layout.
Fig. 12. Cathode–anode differential pressure (DPCA) during the whole test in mGT grid-connected mode.
Fig. 13. Cathode–anode differential pressure (DPCA) in mGT grid-connected mode. The ‘‘max kW’’ load ranges from 74.1 kW at MP = 20 g/s to 71.9 kW at MP = 0 g/s.
component). However, upstream of the OGB (in the stack) the pressure values depend on the pressure losses in the cell sides and on the line performance: while (pcat)in is driven by compressor line performance, (pan)in is mainly a function of fuel mass flow rate and anodic ejector performance. For this reason, the differential pressure on the ceramic material of the stack may be signifi-
cant, especially at part-load conditions and during transient operations. The anodic recirculation loop (vessel and pipes) was designed and built to minimize the cathode–anode differential pressure (DPCA) at design conditions (gas turbine at maximum load and MP = 20 g/s) [28]. The design DPCA value is not null, but it is small (sustainable by the cell materials) as in a typical SOFC system. A DPCA design value under 10 mbar is an acceptable result (the DPCA design value in the test rig is about 7 mbar). Even if a null DPCA design value is desirable, some details in the emulator rig are not exactly evaluated in design calculations (e.g. gaskets, welding details, pipe roughness, bolts, and other installation details), as in a real hybrid plant, and assumptions used may represent real phenomena in a simplified approach. Fig. 12 shows the differential pressure values measured over time during a complete test carried out in electrical grid-connected mode. The initial part (before the ‘‘Test zone’’ of Fig. 12) was necessary to heat the rig pipes and to operate initial MP variations (the DPCA peaks obtained before time 22.40.00). The measurement values obtained after the heating period (inside the ‘‘Test zone’’) were then used to obtain the off-design DPCA trend shown in Fig. 13, required to manage hybrid system operations at part-load conditions. The negative peak at the end of the test is related to gas turbine shutdown phase: since the flow in the ejector primary duct (MP) was maintained at 20 g/s during shutdown, the pressure decrease in the cathodic side (due to the decrease in rotational speed) generates significant negative DPCA values (close to 30 mbar) to be considered for a real hybrid system design. The DPCA value increases with load (for cathodic mass flow rate increase) and with the decrease in MP value. This second trend has to be carefully managed when operating a hybrid system at part-load conditions, to avoid excessive mechanical stress on the cell sides. The maximum sustainable DPCA value depends on fuel cell technology and design. Even if this value for the RRFCS stack is not available for confidentiality reasons, the results obtained with the emulator rig are useful for reducing this stress if considered critical by manufacturer. For this reason, Fig. 13 is very interesting for control strategy development. At part-load condition (MP lower than its design value), the DPCA value can be reduced, decreasing machine load and, as a consequence, the rotational speed (this operation generates a decrease in the cathodic side mass flow rate value). Moreover, an air flow decrease is also necessary to control fuel cell temperature values (see [17] for details), avoiding excessive thermal stress on the materials. The control system must be tuned to manage air mass flow rate and avoid damage due to both mechanical and thermal stress. Tests carried out with emulators are essential to reduce the machine rotational speed at part-load conditions, maintaining both DPCA and fuel cell temperatures under safe conditions. An example of possible dangerous conditions caused by incorrect operations is shown in Fig. 14. A strong step increase in the
376
M.L. Ferrari, A.F. Massardo / Applied Energy 105 (2013) 369–379
Fig. 14. Critical variation of DPCA and compressor outlet pressure (PRC1) due to a surge event generated by MP fast increase (mGT grid-connected mode).
anodic ejector primary duct mass flow rate (from 0 g/s to about 16 g/s) produces a pressure increase in the anodic side (the DPCA value reaches 30.8 mbar after 4 s). In the meantime, this mass flow rate increase step produces an increase oscillation in compressor outlet pressure (PRC1) generating a surge event (detected by the machine vibration sensor with a 3:5 g m/s2 signal). The subsequent cathodic pressure decrease due to the surge cycle produces very dangerous DPCA peak shown in Fig. 14 (minimum value equal to 145.4 mbar), which in a real hybrid system with an SOFC stack would cause the fuel cell to break [35]. 3.2. Stand-alone mode These tests were carried out with different net electrical load values starting from gas turbine idle condition to 60 kW (this is the maximum load to operate tests with MP values up to 20 g/s without excessive stress on the gas turbine fuel compressor). As
for the grid-connected mode, at each electrical load value (idle, 20 kW, 40 kW, 60 kW) different steady-state tests were carried out at various mass flow rate values in the anodic ejector primary duct (MP = 0 g/s, 5 g/s, 10 g/s, 15 g/s, 20 g/s). The case with MP = 0 g/s is justified as a limit condition already discussed for the ‘‘grid-connected’’ tests. Fig. 15 shows the turbine outlet temperature (named TT2 with Fig. 3 nomenclature) and the anodic ejector recirculation ratio (related to its design value [28]) measured operating the gas turbine in stand-alone mode. The turbine outlet temperature significantly increases with load increase because rotational speed is maintained constant in this stand-alone mode. Moreover, the ejector primary duct mass flow increase generates a slight TT2 decrease (at constant load condition). This trend is due to the injection of a pre-compressed mass flow rate, just upstream of combustor inlet, that slightly increases expander power as discussed for the gridconnected mode. In addition, Fig. 15 shows that the recirculation ratio decreases with the load increase (at constant MP value). This is due to the ejector secondary inlet temperature (TEjS1) increase that is generated by the TT2 increase. The TEjS1 variation (from about 556 K at 0 kW load to about 932 K at 60 kW) is significantly higher than the pressure increase at the ejector secondary duct inlet. As shown in [15,28], a temperature increase at the secondary inlet duct of the anodic ejector generates a recirculated mass flow rate decrease (convergent nozzles at subsonic conditions [15]). Fig. 16 shows the ejector pressure rise (DPEj) measured during the tests in stand-alone mode. As already discussed for Fig. 8, this property (DPEj) is almost not significantly dependent by gas turbine load, because (for viscous losses) the pressure increase effect is compensated by temperature variation. Compressor outlet pressure is an important property that is essential to show cathode–anode side interaction related to MP flow injection. Since in this stand-alone mode machine rotational speed is constant, this increase in pressure (PRC1 in Fig. 17) with load increase is not as strong as the increase for grid-connected mode (Fig. 9). It is simply related to compressor pressure/flow map [26] trend at constant rotational speed. Moreover, a slight pressure increase (at constant load condition) with the mass flow rate increase in the ejector primary duct is evident in Fig. 17. In this case, the effect of additional air injection, at constant conditions, is a pressure increase. Since turbine rotational speed is constant (in stand-alone mode), the air injection at combustor inlet generates a slight additional pressure at compressor discharge level. The consequence of this trend is a dangerous decrease in surge margin. Although this aspect is not significant for the turbine in the tests
Fig. 15. Anodic ejector recirculation ratio (F) referred to design value (F0) [28] and mGT turbine outlet temperature (TT2) in stand-alone mode.
M.L. Ferrari, A.F. Massardo / Applied Energy 105 (2013) 369–379
377
Fig. 16. Anodic ejector pressure rise (DPEj) in mGT stand-alone mode. Fig. 19. Cathode–anode differential pressure (DPCA) in mGT stand-alone mode.
there is a decrease in effectiveness for the temperature increase (for electrical load increase) and an opposite behaviour for the test at 60 kW load (as shown in [27]) due to heat transfer coefficient performance. This increase obtained at 60 kW load it is not affecting the general decrease trend with load increase, because the data reported in [27] show that at 72.1 kW (for MP = 0 g/s) the effectiveness is lower than the value obtained for 40 kW load. For these tests, effectiveness is subject to an accuracy ranging from ±6% at 0 kW load to ±3% at 60 kW (see [27] for details). Moreover, at constant electrical load condition, TRC2 shows a slight decrease for TT2 decrease with MP increase. However, this TRC2 decrease is less significant than the TT2 trend and there is a strong e increase effect as MP increases. This is due to an increase in convective heat exchange coefficient on the recuperator hot side obtained with an additional air mass flow rate from the anodic side (as already discussed for the electrical grid-connected mode).
Fig. 17. Compressor outlet pressure (PRC1) in mGT stand-alone mode.
reported here, compressor outlet pressure increase must always be carefully managed for the safe operations of a hybrid system, especially if the turbine has a low surge margin. The recuperator outlet temperature at the cold side (TRC2) and the effectiveness (e) are reported in Fig. 18 for the stand-alone mode tests. TRC2 shows a typical increase trend with load due to TT2 increase. For the effectiveness trend obtained with the increase in the electrical load, the results shown in Fig. 18 are in agreement with the values discussed in [27]. In details, it can be observed that
3.2.1. Cathode–anode differential pressure Also at stand-alone conditions, the cathode–anode differential pressure (Fig. 19) is analyzed to prevent excessive stress on SOFC ceramic material. In this operating mode, the load influence on DPCA is almost negligible because the cathodic mass flow rate variation is not significant, as it is for the grid-connected mode (the rotational speed is constant during all the tests). Moreover, Fig. 19 shows a DPCA increase with the decrease in MP (for each load value), due to the decrease in mass flow rate in the anodic side. As discussed for the grid-connected mode, this aspect has to be considered for the development of control strategies for hybrid systems. The experimental tests reported here show that if used in a hybrid system, a constant rotational speed gas turbine generates further part-load constraints. In details, if rotational speed is constant the cathode–anode differential pressure has to be managed
Fig. 18. Recuperator outlet temperature (TRC2) and effectiveness (e) in mGT stand-alone mode.
378
M.L. Ferrari, A.F. Massardo / Applied Energy 105 (2013) 369–379
through alternative approaches (not always efficient from the turbine point of view). However, if this kind of turbine control strategy is mandatory (e.g., for safety constraints due to stand-alone operations [26]), bleed approaches or mass flow rate bypass solutions (see [24] for details) have to be designed and carefully managed at part-load conditions (to discharge or bypass the excessive amount of compressor outlet flow). The results discussed here show that, if the machine operates with constant speed, specific bleed or bypass strategies are necessary not only to control SOFC temperatures, but also to avoid excessive stress related to high DPCA values.
4. Conclusions This paper presents an experimental analysis of cathode–anode interaction for an SOFC based hybrid system. To avoid the risks and costs related to an actual fuel cell stack, the emulator rig developed at the University of Genoa was used for a broad series of experiments. As demonstrated in previous works [24,25] this kind of test rig (based on the coupling of a microturbine with a complete cathode–anode fuel cell emulator) is able to produce effects similar to real plant results. The following main results were obtained: The recirculation ratio (F) for the anodic ejector showed different trends. While for mGT grid-connected mode F increases when the load increases depending on MP (from about 2% at 5 g/s to about 25% at 20 g/s), in stand-alone mode F decreases with the load increase (almost constant decrement of about 10%). The rise in ejector pressure (DPEj) is not significantly dependent on mGT load for both grid-connected and stand-alone modes. At constant mGT electrical load, for grid-connected mode the increase in anodic ejector mass flow rate in the primary duct (MP) generates a decrease in rotational speed (average variation: 1.4%) while for stand-alone condition it is related to a TOT decrease (variation: 2.6%). At mGT stand-alone mode the air injection on the anodic side (MP) generates an increase in compressor outlet pressure (average variation: <1%) at constant electrical load. This trend is detrimental in terms of the surge margin. Even if it is not significant for these tests, this has to be managed during hybrid system operations (especially with turbines affected by low surge margin performance). From the thermal point of view the trend in recuperator effectiveness showed a slight increase (between 1.1% and 1.7%) for both grid-connected and stand-alone modes (because of the increased heat exchange coefficient due to the additional mass flow rate). Although the performance increase is small, this might be interesting for hybrid system operations. While for mGT grid-connected mode the load increase generates a significant increase in cathode–anode differential pressure (average variation: 12.4 mbar), for mGT stand-alone configuration the electrical load does not appear to significantly influence this property. For both grid-connected and stand-alone modes, cathode– anode differential pressure significantly increases (about 20 mbar) with the decrease in ejector primary duct mass flow. This aspect has to be carefully taken into account to avoid excessive mechanical stress on the fuel cell materials at partload conditions. These experiments are very useful to develop a suitable control strategy to couple the decrease in fuel mass flow rate (emulated with MP) with the decrease in cathode side air.
As discussed in the text this emulator is also equipped with a real-time model [30] for components that are not physically present in the laboratory (e.g. the fuel cell stack). The approach based on hardware–software coupling (see [30] for details) will be essential to develop further tests on cathode–anode interaction during actual operations, focusing attention on critical phases (e.g. startup, shutdown and load changes).
Acknowledgements This work was funded by Felicitas (TIP4-CT-2005-516270) and LARGE-SOFC (No. 019739) European Projects. Additionally, the authors would like to thank Dr. Matteo Pascenti of TPG for his support in the experimental activity. The results presented and discussed here are on the sole responsibility of the authors and do not represent the views of the company cited above.
References [1] International Energy Agency. Key World Energy Statistics, 2006. [2] Kathib H, Energy considerations global warming perspectives, World Energy Council. Energy Permanent Monitoring Panel of the World Federation of Sciences, 2004. [3] Goodstein ES. Economics and the environment. Englewood Cliffs (NJ): Prentice-Hall; 1999. [4] Quirion P. Complying with the kyoto protocol under uncertainty: Taxes or tradable permits? Energy Policy 2010;38(9):5166–73. [5] Veyo S, Shockling LA, Dederer JT, Gillett JE, Lundberg WL. Tubular solid oxide fuel cell/gas turbine hybrid cycle power systems: status. J Eng Gas Turbines Power 2002;124:845–9. [6] Birgersson KE, Balaya P, Chou SK, Yan J. Energy solutions for a sustainable world. Appl Energy 2012;90:1–2. [7] Andersson M, Yuan J, Sundén B. Review on modeling development for multiscale chemical reactions coupled transport phenomena in solid oxide fuel cells. Appl Energy 2010;87:1461–76. [8] Richards GA, McMillian MM, Gemmen RS, Rogers WA, Cully SR. Issues for lowemission, fuel-flexible power systems. Progr Energy Combust Sci 2001;27:141–69. [9] Nanaeda K, Mueller F, Brouwer J, Samuelsen S. Dynamic modeling and evaluation of solid oxide fuel cell – combined heat and power system operating strategies. J Power Sources 2010;195:3176–85. [10] Riensche E, Stimming U, Unverzagt G. Optimization of a 200 kW SOFC cogeneration power plant: Part I: Variation of process parameters. J Power Sources 1998;72:251–6. [11] Mueller F, Brouwer J, Jabbari F, Samuelsen G. Dynamic simulation of an integrated solid oxide fuel cell system including current-based fuel flow control. J Fuel Cell Sci Technol 2006;3:144–55. [12] Calise F, Ferruzzi G, Vanoli L. Parametric exergy analysis of a tubular solid oxide fuel cell (SOFC) stack through finite-volume model. Appl Energy 2009;86:2401–10. [13] Agnew GD, Bozzolo M, Moritz RR, Berenyi S. The design and integration of the Rolls-Royce Fuel Cell Systems 1MW SOFC. ASME Paper GT2005-69122, 2005. [14] Trasino F, Bozzolo M, Magistri L, Massardo AF. Modeling and performance analysis of the Rolls-Royce Fuel Cell systems limited: 1 MW plant. J Eng Gas Turbines Power 2011;133. p. 021701_1–11. [15] Ferrari ML, Pascenti M, Massardo AF. Ejector model for high temperature fuel cell hybrid systems: experimental validation at steady-state and dynamic conditions. J Fuel Cell Sci Technol 2008;5. p. 041005_1–7. [16] Ghigliazza F, Traverso A, Massardo AF, Wingate J, Ferrari ML. Generic real-time modeling of solid oxide fuel cell hybrid systems. J Fuel Cell Sci Technol 2009;6. p. 021312_1–7. [17] Ferrari ML. Solid oxide fuel cell hybrid system: control strategy for stand-alone configurations. J Power Sources, Elsevier 2011;196(5):2682–90. [18] Shibata S, Urata T, Sugawara Y, Umeda T, Morita J, Hatoh K, Kitano Y. Fuel cell system with differential pressure control. U.S. Patent US7691510B2, 2010. [19] Limbeck U, Mazzotta CS, Pasera U, Schmalzriedt S, Wiesner P. Method and apparatus for controlling the differential pressure in a fuel cell. U.S. Patent US2007/0184319A1, 2007. [20] McLarty D, Kuniba Y, Brouwer J, Samuelsen S. Experimental and theoretical evidence for control requirements in solid oxide fuel cell gas turbine hybrid systems. J Power Sources 2012;209:195–203. [21] Massardo AF, Lubelli F. Internal reforming solid oxide fuel cell-gas turbine combined cycles (IRSOFC-GT) – Part I: Cell model and cycle thermodynamic analysis. J Eng Gas Turbines Power 2000;122:27–35. [22] Massardo AF, Magistri L. Internal reforming solid oxide fuel cell gas turbine combined cycles (IRSOFC-GT) – Part II: Exergy and thermoeconomic analyses. J Eng Gas Turbines Power 2003;125:67–74.
M.L. Ferrari, A.F. Massardo / Applied Energy 105 (2013) 369–379 [23] Ferrari ML, Traverso A, Magistri L, Massardo AF. Influence of the anodic recirculation transient behaviour on the SOFC hybrid system performance. J Power Sources 2005;149:22–32. [24] Shelton M, Celik I, Liese E, Tucker D. A study in the process modeling of the startup of fuel cell/gas turbine hybrid systems. J Eng Gas Turbines Power 2010;132. p. 012301_1–8. [25] Hohloch M, Widenhorn A, Lebküchner D, Panne T, Aigner M. Micro gas turbine test rig for hybrid power plant application. ASME Paper GT2008-50443, 2008. [26] Ferrari ML, Pascenti M, Magistri L, Massardo AF. Hybrid system test rig: startup and shutdown physical emulation. J Fuel Cell Sci Technol 2010;7. p. 021005_1–7. [27] Ferrari ML, Pascenti M, Magistri L, Massardo AF. Micro gas turbine recuperator: steady-state and transient experimental investigation. J Eng for Gas Turbines Power 2010;132. p. 022301_1–8. [28] Ferrari ML, Pascenti M, Magistri L, Massardo AF. MGT/HTFC hybrid system emulator test rig: experimental investigation on the anodic recirculation system. J Fuel Cell Sci Technol 2011;8. p. 021012_1–9.
379
[29] Ferrari ML, Pascenti M, Traverso AN, Massardo AF. Hybrid system test rig: chemical composition emulation with steam injection. Appl Energy 2012;97:809–15. [30] Ferrari ML, Pascenti M. Flexible micro gas turbine rig for tests on advanced energy systems, Advance in gas turbine technology. Rijeka, Croatia: INTECH; 2011. ISBN 978-953-308-68-4, p. 89–114, [chapter 5]. [31] Marsano F, Magistri L, Massardo AF. Ejector performance influence on a solid oxide fuel cell anodic recirculation system. J Power Sources 2004;129:216–28. [32] Ferrari ML, Bernardi D, Massardo AF. Design and testing of ejectors for high temperature fuel cell hybrid systems. J Fuel Cell Sci Technol 2006;3:284–91. [33] McDonald CF. Heat recovery exchanger technology for very small gas turbines. Int J Turbo Jet Engines 1996;13:239–61. [34] McDonald CF. Recuperator considerations for future high efficiency microturbines. J Appl Therm Energy 2003;23:1453–87. [35] Ferrari ML, Pascenti M, Bertone R, Magistri L. Hybrid simulation facility based on commercial 100 kWe micro gas turbine. J Fuel Cell Sci Technol 2009;6. p. 031008_1–8.