Modelling and Performance Analysis of the Rolls-Royce Fuel Cell Systems Limited: 1 MW Plant

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MODELLING AND PERFORMANCE ANALYSIS OF THE ROLLS-ROYCE FUEL CELL SYSTEMS LIMITED. 1 MW PLANT

Francesco Trasino Michele BozzoloRolls-Royce Fuel Cell Systems Limited

Loughborough (UK)

Loredana Magistri Aristide F. MassardoThermochemical Power Group (TPG)

Dipartimento di Macchine, Sistemi Energetici e TrasportiUniversità di Genova

Proceedings of ASME Turbo Expo 2009: Power for Land, Sea and Air GT2009

June 8-12, 2009, Orlando, Florida, USA

GT2009-59328

ABSTRACTThis paper is focused on the performance of the 1MW plant designed and developed by Rolls-Royce Fuel Cell Systems Limited. The system consists of a two stage turbogenerator coupled with pressure vessels containing the fuel cell stack, internal reformer, cathode ejector, anode ejector and off gas burner. While the overall scheme is relatively simple, due to the limited number of components, the interaction between the components is complex and the system behaviour is determined by many parameters. In particular two important subsystems such as the cathode and the anode recycle loops must be carefully analyzed also considering their interaction with and influence on the turbogenerator performance. The system performance model represents the whole and each physical component is modelled in detail as a sub-system. The component models have been validated or are under verification. The model provides all the operating parameters in each characteristic point of the plant and a complete distribution of thermodynamics and chemical parameters inside the SOFC stack and reformer. In order to characterise the system behaviour, its operating envelope has been calculated taking into account the effect of ambient temperature and pressure as described in the paper. Given the complexity of the system various constraints have to be considered in order to obtain a safe operating condition not only for the system as a whole but also for each of its parts. In particular each point calculated has to comply with several constraints such as stack temperature distribution, maximum and minimum temperatures and high and low pressure spool maximum rotational speeds. The model developed and the results presented in the paper provide important information for the definition of an appropriate control strategy and a first step in the development of a robust and optimized control system.

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INTRODUCTIONThe integrated SOFC-GT hybrid system layout considered in this work is the Rolls-Royce Fuel Cell Systems Ltd (RRFCS) 1MW System layout; in particular since the system is a modular one made up of four 250kW units called Generator Module (GM), the actual system layout studied is the one of a single GM reported in Figure 1. Each GM is a virtual stand-alone unit capable of operating on its own and sharing a limited number of auxiliaries with the other GM part of the 1MW package. Each GM consists of a two stage turbogenerator coupled with pressure vessels containing sub-systems called tiers; tiers contain the fuel cell stack, internal reformer, cathode ejector, anode ejector and off-gas burner. Figure 1 shows the general cycle arrangement and the main components: while the overall scheme is relatively simple due to the limited number of components, the interaction between components is complex and the system behaviour is determined by many parameters.Two important sub-systems operate within this system:

• the cathode (air) recycle (Figure 2),• the anode (fuel) recycle (Figure 3).

The fresh air flow from the system inlet is pre-heated by recycling a fraction of the cathode side exhaust. The resulting mixture, at ~850ºC, feeds the stack. The recycle is driven by an ejector where higher pressure primary air entrains hot secondary cathode exhaust. The recycle ratio (secondary mass flow / primary mass flow), for a given ejector design, is a function of the pressure and temperature of the primary flow, the temperature of the secondary flow, and the pressure loss to be overcome in the recycle loop. Pressure losses are functions of the mass flows, temperatures and compositions of the various streams along the loop.

1 Copyright © 2009 by ASMECopyright © 2009 Rolls-Royce Fuel Cell Systems Limited

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Figure 1: RRFCS system cycle and key components.

850°C 7 ATM

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Figure 2: Cathode recycle loop.

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Figure 3: Anode recycle loop.

On the anode side, a fraction of the steam rich stack off-gas is recycled into the incoming methane based fuel to pre-heat the fuel and to recycle enough steam for the methane reforming reaction and moreover to disadvantage carbon deposition; this event should be avoided since carbon greatly reduces catalyst activity hence its deposition on active sites in both reformer and stack would be detrimental for system performance and safe operation. The methane / steam mixture is then converted in a reforming rector into a hydrogen rich mixture that is used as the anode side reactant in the fuel cell. The recycle is driven also in this case by an ejector where primary fuel entrains the steam rich secondary fuel from stack outlet. Excess fuel discharged by the anode loop feeds the off-gas burner where it burns with air in the cathode recycle loop to help keep the stack in temperature.

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The turbogenerator is, as mentioned above, a two stage machine where the low pressure shaft is connected to an alternator while the high pressure shaft is free rotating. Contrary to standard two stage machines it is designed to operate in an unbalanced way since high pressure turbine inlet is at system pressure, around 7 bar, while the high pressure compressor outlet, i.e. cathodic ejector primary, is at a higher pressure due to the need of driving the recycling in the ejector. A blow off valve is mounted at the high pressure compressor outlet to allow the recycling of part of the flow at the low pressure compressor inlet at part load and to avoid surge of the high pressure compressor itself or over speed of the high pressure spool.Each GM is coupled with a fuel pre-processor to provide a suitable fuel for the system: its main purposes are desulphurise the natural gas from the grid and to provide specific gas mixtures for start-up, shut-down and emergency procedures.In addition to the Generator Module components there are two other primary systems included in the 1MW package:

• An electrical power system required to invert the DC power to AC power, condition it, supply power to the plant auxiliaries, and provide a connection to the electricity utility grid.

• A control system required to monitor and sequence operations, and to ensure that all components operate safely and reliably.

STEADY STATE SYSTEM MODELThe system performance model was developed in Matlab-Simulink® environment and comprises different sub-models that were designed from scratch or upgraded from previous ones and tested or validated were possible; in particular it includes: ejector steady state models for both cathode and anode loop, cathodic and anodic loop circuit characteristic models, detailed reformer model, detailed stack model, Low and High Pressure Compressor models (LPC, HPC), Low and High Pressure Turbine models (HPT, LPT), off-gas burner model; non negligible heat exchanges between certain components or with the ambient are also modelled. The model calculates all main parameters such as pressure, temperature, mass flow and composition at inlet and outlet of all the components; a list of all the main point is reported in Table 1 and Table 2.

0a Natural gas to 1MW system1a Natural gas to EFP2a1 Safe gas to anodic ejector primary2a Natural gas from EFP3a Anodic ejector primary from EFP4a1 Anodic ejector outlet4a Reformer inlet5a Reformer outlet6a Stack outlet7a2 Additional syngas to OGB7a1 Stack outlet to OGB7a Stack outlet to anodic ejector secondary inlet

Table 1: List of main calculation point and section in system model for the anodic loop.


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0c1 Recirculated flow from bleed 0c Fresh air from ambient 1c LP compressor inlet 2c1 LP compressor outlet 2c HP compressor inlet 3c1 HP compressor outlet

3cHP compressor outlet after heat exchange with HP turbine outlet

4cCathodic ejector primary after heat exchange with ejector secondary

5c1 Cathodic ejector outlet 5c Cathodic ejector outlet after heat loss to ambient 6c Stack outlet 7c1 Reformer outlet 7c Reformer outlet after heat loss to ambient 8c2 Reformer outlet to OGB 8c1 Reformer outlet to HPT 8c HP turbine inlet after heat loss to ambient 9c1 HP turbine outlet

9cLP inlet after heat exchange with HP compressor outlet

10c LP turbine outlet to ambient 11c1 OGB outlet to cathodic ejector secondary inlet

11cCathodic ejector secondary after heat exchange with primary

Table 2: List of main calculation point and section in system model for the cathodic loop.

Together with these information, the model provide also a complete distribution of the above parameters inside stack and reformer. A brief description of each module of the model is reported in the following

TurbomachineryThe turbogenerator model includes an interpolator for mass flow, pressure ratio and efficiency as function of the rotating speed for the compressors and turbines; the characteristic maps interpolated are provided by the manufacturer [1]: all curves are referred to design conditions and can be easily updated. As already said, the turbogenerator is a two stage one where the low pressure spool is controlled by the alternator while the high pressure one is free to rotate and adjust its speed as required by the operating condition of the machine: LP spool rotational speed is an input for both turbine and compressor (hence alternator power is an output), while HP spool rotational speed is interpolated to give null power output minus losses. The model implements also a model for the bleed valve installed at the outlet of the HP compressor; this valve is used to bypass part of the outlet mass flow at the inlet of the LP compressor to avoid approaching the surge margin of the HP compressor and to avoid possible over speed of the HP spool. In this bleed module the user can define a percentage of the HPC outlet mass flow that is blown off and at this point two operating mode are possible: full recycle mode where all the blown off mass flow is bypassed at LPC inlet or discrete recycle mode where only the mass flow passing through a fixed choked orifice (whose dimension is used defined) is recycled and the remaining fraction is considered blown off to ambient. The model includes also a number of operating information of the real turbogenerator such as inlet, outlet and between

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turbomachinery pressure drops as well as leakages if presents and heat exchanges.

EjectorsEjectors were modelled using a 0-D approach considering only the main sections of the ejector, as reported in Figure 4 where 1 is nozzle section, 2 is secondary section, 3 is mixing duct section and 4 is diffuser outlet section; the model can also take into account the thickness of the nozzle by using fluid dynamic sections obtained by correction of the geometrical ones as shown in Figure 4.

Figure 4: Ejector main sections and detail of nozzle sections.

The model implements energy, continuity and momentum balance that in this case can be simplified as follow:

• Energy equation: m3∗h

t3=m

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Figure 5: Ejector performance compared to experimental results: (m) and (c) stand for measured and calculated respectively.

Anodic and cathodic circuitsAnodic and cathodic loops were modelled considering that in each section of the circuit (stack, reformer, off-gas burner, pipes) pressure losses were predominantly of a certain type. The pressure losses along both circuits are then calculated at the inlet of each reference section using the Darcy equation:

p=0.5∗ f ∗2∗∗l /Dwhere by using the appropriate expression for f in case of laminar, turbulent or pure turbulent flow, pressure drops can be expressed through the following:

p=x∗m y∗R∗T / p∗ p∗p /x∗my∗R∗T ref

where x and y vary with the type of flow considered in the section (i.e. laminar turbulent or pure turbulent) and subscript ref stands for reference conditions; pressure drops are thus expressed as a function of temperature, composition, mass flow and viscosity and reference design conditions for each main circuit section. Both anode and cathode loop circuit models have been validated against the experimental characteristic curves for reference design condition.

ReformerThe reformer is modelled as an adiabatic co-flow metallic heat exchanger through a Matlab .dll; it’s a detailed multi node model where heat exchange areas and internal geometry can be defined as required by changing the number of nodes. For each node the model calculates pressure, temperatures, flow and composition. The model considers methane as well as higher hydrocarbon species reforming and water gas shift reactions using an equilibrium approach method, no kinetics has been and will be included:

• Methane reforming: CH4 + H2O = CO + 3H2

• Water gas shift: CO + H2O = CO2 + H2

The model was verified against a test case, and was finally updated with the thermo-physical properties and geometry of the reformer installed into the system [2].

StackRRFCS stack design is based on a multi cell pattern screen printed onto a flat ceramic tube; cells are printed on both sides

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of the ceramic tube with fuel flowing in the internal channels, air flowing on the outside crosswise: tubes are then packed to form bundles and bundles are packed vertically to create strips. Strips are assembled to form block that are then packed together in tiers; tiers are enclosed into the pressure vessel (Figure 6). From the flow point of view (Figure 7), tubes in a bundle are in series on the anodic side whit the cathodic stream flowing cross-flow wise to the tubes; bundles in a strip are in parallel on both anodic and cathodic side. Strips are in series on the cathodic side while in parallel on the anodic one. From the electrical point of view, tubes and bundle are in series, strips are in parallel.

StripBundleBlock

StripBundleBlock

StripBundleBlock

Figure 6: A block of RRFCS strips.

Figure 7: Flow scheme for a stack block.

Only a portion of the stack corresponding to a single tier of a block (highlighted in green in Figure 7) was actually modelled and replicated to match the exact amount of stack in the system; this decision was based on the following assumptions:

1. Uniform distribution of the anodic flow between bundles.

2. Uniform distribution of anodic flow and temperature at bundle inlet.

3. Uniform distribution of the cathodic flow between blocks.

4. Uniform distribution of cathodic flow and temperature at block inlet.

Given the modularity of the SOFC stack, this is the smallest portion of it that allows to fully replicate the flow scheme of a block and at the same time reduces computational time to an



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acceptable level. These assumption despite being somewhat constraining are reasonable anyway since a sound system design should provide an as much as possible uniform distribution of flows; as a consequence of these assumption, a single tier of a block is thus considered adiabatic.The block single tier model was developed starting from an already existing SOFC model developed in Fortran language [3]; the code was upgraded to represent the RRFCS planar design and converted to a Fortran Mex-Function for use with Matlab-Simulink®; the model thus modified has the following features and underlying assumption:

1. Each tube is divided in n cell corresponding to the actual number of cells printed on a real tube: each cell is considered as a wafer of porous ceramic tube, anode, electrolyte, cathode.

2. Real geometrical and electrical data: thickness of anode, cathode, electrolyte current collectors, cell active area.

3. Voltages and losses (ohmic, contact, activation and concentration polarisation) calculated for each cell.

4. Current is uniform for all the cells and for both sides of the printed tube for a complete symmetry between upper and lower halves of each tube.

5. Heat exchange between adjacent cells and between cell and anodic and cathodic flows is calculated.

6. Each bundle is considered adiabatic.7. Reforming and water gas shift reactions considered at

equilibrium for each cell.This model was finally upgraded to a C Mex-function of a single bundle thus allowing the following improvements:

1. Dynamic sizing of the n° of cells per tube and n° of tubes per bundle thus avoid the need of recompiling the model each time these parameters have to be changed to adapt to a new tube and stack design.

2. Reduced calculation time through the use of discrete states allowing a faster convergence in calculations.

3. Heat exchange between facing tubes.4. Current of each bundle is modified in order to

maintain the imposed average operating current and to provide a uniform voltage distribution: this allows to mimic the electrical parallel connection of the strips, hence of bundles, in the real system.

SYSTEM PERFORMANCEIn order to characterise the system behaviour, its operating envelope was calculated taking into account the effect of ambient temperature and pressure. Given the complexity of the system, various constraints have to be considered in order to obtain a suitable operating condition not only for the system as a whole but also for each of its part; in particular each point calculated had to comply with limits on stack temperature distribution, maximum and minimum temperatures, limits on high and low pressure spool rotational speed and on the depressurisation of the system. The system model used for the calculations integrated the detailed stack model since only in this way hot or cold point or unfeasible temperature distributions inside the stack could be detected as a result of the system operating condition. A drawback of this approach was an increase in calculation time given the detail of the stack


model but it was deemed acceptable in view of the information required. The design point of the system is the result of the minimum target performances set by RRFCS for the first generation GM: 250 kW net power output at 56% net efficiency in design conditions (i.e. ISO conditions of ambient temperature and pressure, 15 °C and 1.013 bar ), including all auxiliaries and losses related to a single GM.

Part load operationSystem performance is mainly determined by the stack one since the main function of the turbogenerator is to pressurise the system and its contribution to power output is very limited: hence each operating point is mainly determined by reducing or increasing the current drawn hence reducing or increasing the power output of the stack. The first operating envelopes were calculated with the aim of understanding the behaviour of the system and its response to changes in control parameters; a final one was later calculated drawing from the experience of the first ones with the aim of optimising system performance also in view of the definition of a possible control system. All operating points were calculated reducing the current from the design value down to 50% and varying fuel and air flow at the same time to comply with the following constraints for a safe system operation:

1. Stack temperature cannot exceed maximum and, to a lesser degree, minimum temperature limits; maximum temperature is far below usual operating temperature for SOFC found in literature (1000-1100 °C) since operating life decreases with the increase in temperature and a reduced operating life cannot be accepted in a base load power generation system.

2. Surge condition for both low and high pressure compressor must be avoided: hence a minimum safe limit was always maintained.

3. High pressure spool over speed must be avoided: this condition could happen at part load, where the low pressure spool speed is reduced through the alternator to decrease the air flow and at the same time the high pressure spool speed tend to increase since the inlet temperature of the high pressure turbine remains almost constant.

The 0-50% operating range is not part of the study as continuous operation in this part of the envelope is considered unlikely. In particular, benefiting from the modular architecture any power output can be obtained with an appropriate combination of working generator modules.

Part load operation: ambient temperatureOn a general basis, ambient parameters that can affect performance of SOFC hybrid systems are temperature, pressure and air humidity. Ambient temperature and pressure are by far the most important since they greatly affect turbomachinery performance and hence indirectly affect the performance of pressurised hybrid systems relying on them. The effect of pressure, in general of a decrease in pressure related to altitude, is a reduction in turbomachinery performance in terms of power output and moreover in terms of air-flow delivered by the compressor. In this paragraph the effect of ambient temperature on system performance will be the focus of the analysis.



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Various operating curves for different ambient temperature have been simulated; the load has been reduced from 100% to 50% nominal load by reducing the operating current of the fuel cell itself. A reduction in operating current causes a reduction in the heat released by the operation of the stack, causing as a consequence a cooling of the stack itself if no action were to be taken. In order to ensure a proper thermal management of the system and consequently avoid any relevant changes in stack temperature, two parameters, fuel and air mass-flow, can be effectively controlled:

• Rotational speed of low pressure compressor: air flow is primarily reduced by reducing the rotational speed of the low pressure compressor through the alternator (alternator control).

• Bleed valve: a reduction in fresh air flow to the cathodic primary can be obtained by blowing off part of the air flow at the outlet of the high pressure compressor through the bleed valve (bleed control).

• Fuel flow: reducing or increasing the fuel flow, hence increasing or reducing the fuel utilisation factor (Uf), varies the fraction of unspent fuel in the anodic off-gas burning in the off-gas burner with part of the cathodic flow thus increasing the overall system temperature through the cathodic recirculation (Uf control).

A reduction in air flow is primarily obtained by reducing the rotational speed of the turbomachinery that also causes a decrease in system operating pressure; blowing-off part of the air at the high pressure compressor outlet through the bleed valve yields the same result and although being helpful in avoiding surge conditions of the high pressure compressor itself is obviously less efficient since part of the air is discharged to the ambient without doing any useful work. Also in this case system operative pressure decreases although comparatively less than the previous case due to the change in the turbomachinery operating condition. Obviously in case of an increase in the operating current, both parameters could be effectively controlled in the opposite way to avoid an increase in system temperature. From these preliminary considerations, the effect of ambient temperature could be considered, at least in principle, similar to an increase or reduction in current since an increase or reduction in ambient temperature from the standard reference temperature of 288.15 K should cause a corresponding increase or reduction in stack temperature by means of the cathodic recirculation.Figure 8, Figure 9, Figure 10, Figure 11 show respectively part load performance, stack DC power, Uf and operating pressure as function of ambient temperature. The part-load curve at +15ºC (288.15 K) is the reference one as already said: for this curve only, the design point of the system is reported together with the operating point at 110% of the design operating current. Since in general a part of the system power output is due to the turbomachinery alternator, the system power reduction is not directly proportional to the current reduction but tends to be smaller as seen from Figure 8.As shown in Figure 8, part-load operation is favoured with high ambient temperature while the opposite happens in near design point condition. This behaviour can be explained by considering a generic gas turbine behaviour as function of temperature: at part-load (Figure 8, A area) a higher ambient


temperature means a higher temperature at the compressor outlet and hence more heat to the system through the cathodic recirculation with lesser need to reduce the air-flow (or Uf, Figure 10) to avoid stack cooling: as a consequence the alternator delivers a little more power, the system operating pressure remains higher (Figure 11), stack operates more efficiently since a higher Uf means fuel is converted more efficiently through the electrochemical reaction hence the overall performance is better than with a lower ambient temperature. Whereas near full-load (Figure 8, B area), system performance is deeply affected by turbomachinery behaviour: a higher ambient temperature means the compressor will deliver less air hence the need to raise the rotational speed to increase the air flow and avoid over-heating the system; the system pressure increases but the power delivered by the alternator decreases with the increase in ambient temperature due to the greater power demand to drive the compressor hence the overall system performance is worse than with a lower ambient temperature. Obviously the rotational speed can be increased only to a certain limit for mechanical reasons, after that the only means to avoid overheating the system are reducing the operating current or increasing the Uf; reducing the current implies reducing the fuel cell power output whereas Uf can be increased up to usually no more than 0.85 to ensure sufficient margin from localised fuel starvation. This last occurrence is more likely to happen during very hot days or while operating at higher current as can be seen from Figure 10 at +30ºC 100% current as well as +15 ºC 110% current were Uf has to be increased to manage stack heat balance. The need to decrease or increase the air-flow in order to keep the fuel cell stack temperature in a safe range causes also a change in the operating condition of the turbomachinery that ultimately causes respectively a decrease or an increase in the system operating pressure as can be seen from Figure 11, where the arrow shows the increase in both air-flow and turbomachinery rotational speed. To this regard it should be noted that at the onset of the study pressure cycling in the system was considered unsafe hence system depressurisation was limited (Figure 11): for this reason and to avoid surge of the high pressure compressor at part load, the blown-off fraction of the air at high pressure compressor outlet increases with the reduction of load and temperature. Considering both Figure 8 and Figure 10, at higher ambient temperature the system can work at a higher Uf for the same operating current and for a wider range of part load conditions thus in general performing better than at low ambient temperature. This effect is clearly visible considering the +30ºC where in the range around 80÷90% load efficiency is almost constant; obviously this effect is greater the higher the ambient temperature. As it will be shown later, the constant Uf control strategy at part-load operation is the most effective from the efficiency point of view.The effects of ambient temperature can be summarised as follow:

1. Part-load performance increases with ambient temperature since a high temperature helps in keeping the system in temperature.

2. Full-load performance is favoured by lower ambient temperatures, at least in the range +30ºC ÷ -20ºC



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where system performance is deeply affected by turbomachinery behaviour.

3. Issues on thermal management of the system may arise in conditions such as very hot day and full-load when the system might over-heat due to the reduced air-flow delivered by the compressor.

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Figure 11: Pressure trend at cathodic ejector outlet i.e. stack inlet at cathodic side.

Part load operation: ambient pressureThe behaviour of the system in view of a reduction in ambient pressure such as in case of installation not on sea level is the focus of this paragraph: the effect of the standard fluctuation in ambient pressure at sea level or at a given altitude due can be easily inferred from the results here reported as a particular case with a more limited effect. As briefly stated before ambient pressure has a twofold effect on system performance:

• It reduces turbomachinery performance by decreasing the air-flow delivered by the compressor and as a consequence the alternator power output: a reduction in power output can be more or less relevant depending on the power share demanded to the turbomachinery; in case of the RRFCS design the turbogenerator has the main function to pressurise the system and not provide a relevant fraction of the overall power so this effect is somewhat mitigated.

• It diminishes stack performance: a reduction in air-flow (see previous point) means a decrease in coolant in the system for a given speed of the turbogenerator hence the stack would heat-up unless air-flow is increased as required up to the maximum limit allowed by the compressor or the heat generated by the fuel cell itself is reduced by lowering the operating current or the fuel-flow.

Figure 12 show a comparison between the operating curves at ambient temperature 288.15 K at 0m and 2000m above mean sea level (amsl); Figure 13, Figure 14 and Figure 15 show the same parameters as reported in the previous paragraph considering the system as if operating at an altitude of 2000m: operating curves were calculated for only a couple representative curves since as expected the behaviour as a function of the temperature is quite similar to the ambient pressure case; values are again referred to full load ISO condition except where stated. Figure 12 clearly shows that system performance could be limited to less than 85% of the power at ISO condition at sea level being the actual current drawn 86% of nominal value: this can be easily explained recalling that in this case the air-flow is limited by the maximum value the turbogenerator can deliver without over-speeding hence operating current is limited to the maximum value allowing a safe temperature distribution inside the fuel



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cell stack. As could be expected considering turbomachinery behaviour, at lower ambient temperature i.e. at lower compressor inlet temperature, a higher current can be drawn (Figure 13) since higher is the air flow delivered. The high values of system efficiency at reduced part load can be explained considering that in this case the reduced air-flow delivered is somewhat beneficial to the thermal management of the stack allowing to maintain an almost nominal Uf for a wider range compared to the ambient pressure case (Figure 14) and at the same time greatly reduce the use of the bleed valve thus reducing its detrimental effect on system efficiency. Figure 15 finally shows stack inlet pressure at the anodic side: its is obviously similar to the stack inlet pressure at the cathodic side shown in Figure 11 given the limit on differential pressure between anodic and cathodic side.

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Figure 12: Operating curve comparison for ISO temperature.

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Figure 13: GM efficiency at 2000m amsl.

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Figure 14: Fuel utilisation factor at 2000m amsl.

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Figure 15: Pressure at stack inlet, anodic side, at 2000m amsl.

Control strategy definitionThe available data reported in part in the operating envelopes presented (Figure 8) were used to understand the behaviour of the system and its response to changes of different control parameters and of a combination of them. They were subsequently used as a starting point in the optimisation of the part load control strategy of the system. From the calculations carried out, two main part load control strategies could be highlighted:

1. GT & Uf (1): main control parameter is Uf, bleed valve kept almost fixed after initial opening, limited reduction of low pressure spool speed.

2. GT & Uf (2): Uf and bleed valve are main control parameters continuously varying along the operating curve, limited reduction of low pressure spool speed but higher than previous case.

A third control strategy can be inferred from the results of the previous two as the optimum one but only in case where there is no limit on the depressurisation of the system to avoid pressure cycling:

3. GT & Uf (3): low pressure spool speed and bleed are controlled continuously varying along the operating curve, Uf kept at nominal value.

Figure 16, Figure 17, Figure 18 and Figure 19 show part load performances and operating parameters for the above mentioned control strategies; all data is referred to the design point condition. Figure 18 clearly shows that strategy 2 and even better strategy 1, despite limiting the reduction in system operating pressure and hence limiting the reduction of fresh air flow delivered by the turbogenerator, (see Figure 17 showing low pressure compressor rotational speed), heavily rely on Uf

and bleed fraction to ensure a proper thermal management of the system. As can be seen from Figure 16, strategy 1 is the worst performing while 3 is the best one in terms of efficiency but since it relies most on reduction of the turbomachinery rotational speed to manage thermal balance of the stack it causes the greatest reduction in operating pressure (Figure 18). Strategy 1 on the contrary is at the opposite since it relies mostly based on Uf reduction to manage thermal balance hence the reduction in system pressure is limited; strategy 2 is a compromise between the previous two. In general it can be said that reducing Uf directly affects the efficiency of the stack and since this makes up for most of the system efficiency, Uf


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reduction should be avoided. Obviously there could be a limit on how much pressure reduction could be accepted both in order to avoid pressure cycling or because of limits on the minimum rotational speed of the turbomachinery and both these requirement could limit the extent of the constant Uf

operation; moreover the design of the control system could impose limits on the kind of part-load control strategy that can be applied. As a general indication in part load operation a constant Uf strategy should be applied as long as possible; if and when this condition could not be applied, Uf decrease should be nonetheless reduced to a minimum.

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Figure 16: Overall system performance.

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Figure 17: Low pressure compressor rotational speed.

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Figure 19: Fuel utilisation factor.

Blow-off recycleAs already noted, part of the air at high pressure compressor outlet can be blown-off to ambient through the bleed valve to avoid surge of the high pressure compressor. During the definition of the optimum part load control strategy, the possibility of recycling part or all of the blown-off air at the low pressure compressor inlet was considered: blown-off fractions becomes non negligible with load reduction and low ambient temperature decrease thus greatly affecting overall system efficiency. Recycling the blown-off air would increase air temperature at low pressure compressor inlet thus de-facto changing the operating point of the turbogenerator that would see an apparently higher “ambient temperature”. As with high ambient temperature, this change in operating point would cause a reduction in air flow delivered by the compressors and an increase of compressor outlet temperature, both beneficial effects for the thermal management of the stack at reduced load. As a consequence, operating curves moves upward, with operating points moving towards operating points at an equivalent compressor inlet temperature. This can be clearly seen from Figure 20 where the operating curve at ambient temperature 223.15 K (dark blue solid line) moves upward (dark blue dashed line) starting from around 100% design current at ambient temperature 253.15 K and ending at around 70% design current at ambient temperature 273.15 K. For these two operating points in fact low pressure compressor inlet temperature is 260 K and 269 K respectively and other turbogenerator parameters (like rotational speed, mass flows, pressures, blow-off fractions) are in between of those of the two operating points without blow-off recycle.

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Figure 20: Effects of blown-off air recycling at compressor inlet on system performance.

Conclusions: Operating envelope optimisationFigure 21 to Figure 26 show part of the operating envelope of the system and some of the operating parameters that were simulated with the aim of optimizing the operating envelope in terms of efficiency in order to provide useful information for the definition of a reliable control system and with the following assumption:

1. Fuel utilisation factor has been kept constant since this control strategy was previously highlighted as the optimum in terms of overall system performance.

2. Air flow is solely controlled through low pressure compressor rotational speed.

3. Use of bleed valve is limited to avoid surge of the high pressure compressor and to avoid over speed of the high pressure spool: all the operating points are calculated at minimum surge margin for the high pressure compressor.

4. Blown off air at high pressure compressor outlet is fully recirculated at low pressure compressor inlet.

5. No calculations were done at load less than 50% (i.e. system current less than 50% of the nominal full load value) for the reasons already reported (see “Part loadoperation”).

As shown in Figure 21, system overall behaviour is quite similar to the one previously reported (Figure 8), presenting a part load area (A) where performance is mainly influenced by the effect of ambient temperature on fuel cell performance and an almost nominal load area (B) where system performance is mainly influenced by the effect of ambient temperature on turbomachinery behaviour. It is clearly evident the non negligible increase in GM efficiency over the whole operating range and for all temperatures that remains above 95% of the nominal value compared to an 80% of the previous case: this is due primarily to the use of a constant Uf and in part to the recycle of blown-off air at part load (see Figure 20).Contrary to previous calculations, the turbogenerator act almost solely as a turbocharger since its power share is at most 5%, while stack power is almost proportional with the current (Figure 22). The non proportionality between GM power and current is even more evident in this case where the turbogenerator contribution is negligible and not enough to compensate losses and part of the reduction in performance due to system depressurisation (Figure 23); the effect of system depressurisation is even more evident since average cell voltage decreases despite reducing the operating current (Figure 24).The reduction in system pressure is due to a change in operating point of the cathodic ejector (see Figure 25, Cathodic recirculation factor) caused by a reduction of pressure and fresh air at the high pressure compressor outlet (i.e. cathodic ejector primary, see Figure 23 for fresh air trend since they are similar). Cathodic air has a twofold effect: it acts as coolant (or the opposite depending on the heat released at the off-gas burner) and it helps smooth temperature gradients inside the stack, in particular from first bundle to last one. As a consequence of the reduction in cathodic air, temperatures gradients between inlet and outlet of the stack tend to increase (Figure 26): hence,


given the limit on the maximum stack temperature that is usually reached at the outlet (see asymptotic trend of outlet temperature at reduced load), the inlet temperature decreases with the reduction in load and with temperature until a point when it’s not acceptable anymore (dotted red line). In this case, a reduction in Uf could be beneficial as shown by the points on Figure 26 despite reducing global GM efficiency (Figure 21).

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Figure 21: Single Generator Module operating envelope.

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Figure 22: SOFC stack DC power.

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Figure 23: System operating pressure.



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Figure 24: Average operating cell voltage.

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Figure 25: Cathodic recirculation factor.

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Figure 26: Average stack inlet and outlet temperature (values are referred to inlet stack temperature in ISO conditions).

NOMENCLATURE

RRFCS Rolls-Royce Fuel Cell Systems LimitedGM Generator ModuleGT Gas TurbineLPC Low Pressure CompressorsHPC High Pressure CompressorsHPT High Pressure TurbineLPT Low Pressure TurbineUf Fuel utilisation factor

loaded From: http://proceedings.asmedigitalcollection.asme.org/ on 10/13/2014 Ter

mi mass flow at ejector section ihi enthalpy at ejector section iMi Mach number at ejector section I�i section i of ejector�i cp/cv ratiov average velocity� viscosityf friction factorl characteristic lengthD characteristic diameter� densityR gas constantp pressuredp pressure dropT temperature

ACKNOWLEDGMENTSPut acknowledgments here.

REFERENCES[1] Rolls-Royce Corporation.[2] D. Bernardi, 2007, “Design and analysis of the anode cycle in a SOFC Hybrid system”, PhD Thesis, University of Genoa, DiMSeT.[3] P. Costamagna, A. Selimovic, M. Del Borghi, G. Agnew, 2004, “Electrochemical model of the integrated planar solid oxide fuel cell (IP-SOFC)”, Chemical Engineering Journal 102 (2004) pp 61-69.[4] L. Magistri, M. Bozzolo, O. Tarnowski, G. Agnew A.F. Massardo, 2003, “Design and Off-Design Analysis of a MW Hybrid System Based on Rolls-Royce Integrated Planar SOFC”, ASME Paper GT-2003-38220.[5] M. Ferrari, M. Pascenti, A. F. Massardo, 2008“Ejector Model for High Temperature Fuel Cell Hybrid Systems: Experimental Validation at Steady-State and Dynamic Conditions”, Journal of Fuel Cell Science and Technology, Vol. 5, 041005, 1-7[6] L. Magistri, A. Traverso, F. Cerutti, M. Bozzolo, P. Costamagna, A. F. Massardo, 2005, “Modelling of Pressurised Hybrid Systems Based on Integrated Planar Solid Oxide Fuel Cell (IP-SOFC) Technology”, Fuel Cells, Topical Issue “Modelling of Fuel Cell Systems”, Vol. 1, Issue 5, February 2005, WILEY-VCH.[7] F. Trasino, 2008, “Fuel flexibility in SOFC based hybrid systems”, PhD Thesis, University of Genoa, DiMSeT.



Modelling and Performance Analysis of the Rolls-Royce Fuel Cell Systems Limited: 1 MW Plant

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Transcript of Modelling and Performance Analysis of the Rolls-Royce Fuel Cell Systems Limited: 1 MW Plant