MCFC-based marine APU: Comparison between conventional ATR and cracking coupled with SR integrated inside the stack pressurized vessel

international journal of hydrogen energy 34 (2009) 2026–2042

Available at www.sciencedirect.com

journal homepage: www.elsevier.com/locate/he

MCFC-based marine APU: Comparison between conventional ATR and cracking coupled with SR integrated inside the stack pressurized vessel S. Bensaida, S. Specchiaa,*, F. Federicib, G. Saraccoa, V. Specchiaa a

Materials Science and Chemical Engineering Department, Politecnico di Torino, Corso Duca degli Abruzzi, 24 - 10129 Torino, Italy Ansaldo Fuel Cells S.p.A., Corso Perrone 25, 16152 Genova, Italy

b

article info

abstract

Article history:

In the present work the implementation of MCFCs as auxiliary power units on-board large

Received 19 August 2008

vessels, such as cruising, passengers or commercial, ships was investigated. The MCFC

Received in revised form

stack was designed to supply 500 kWe and was fed with diesel oil undergoing a reforming

23 October 2008

process. The system modelling of the plant was performed in steady-state and aimed at

Accepted 29 November 2008

assessing the power efficiency for different reforming strategies, process configurations

Available online 6 January 2009

and constituting items thermal integrations. The code Matlab/Simulink was used to this end. Two major fuel processing strategies were examined: ‘‘auto-thermal reforming’’ and

Keywords:

‘‘inside vessel steam reforming’’. The latter consisted of a pre-reforming unit in which the

Molten carbonate fuel cell

liquid fuel underwent a catalytic cracking in mild conditions; subsequently, the resulting

APU system

gas mixture made of light hydrocarbons was mixed with steam and fed into a steam

Naval applications

reformer inside the MCFC stack vessel, where conversion to syngas occurred. Due to the

System modelling

high temperature (650  C) exothermic level of MCFC, the stack was compatible with

Reforming

a syngas steam reforming production thermally self sustained. This allowed to increase the

Thermal integration

global electrical efficiency from 32.7% (for the ATR-based system) up to 44.6%. The process

CHP system

was then designed aiming at increasing the overall efficiency by thermally integrating the outlet flue gases with the pre-heating section. This lead to efficiencies equal to 39.1% and 50.6% for the ‘‘auto-thermal reforming’’ and ‘‘inside vessel steam reforming’’, respectively. Finally, the process was upgraded from an auxiliary power unit (APU) to a combined heat and power unit (CHP), since the residual heat in the flue gases was recovered for heating purposes (sanitary water production) and the demineralised water recirculation was implemented to reduce the water make-up and the process environmental footprint. ª 2008 International Association for Hydrogen Energy. Published by Elsevier Ltd. All rights reserved.

1.

Introduction

A clear trend towards the design and installation of integrated electric propulsion systems in ships has emerged in the last few years [1–3]. Most of the cruise ships employ diesel engines

to produce propulsion and diesel based generators for the hotel power for ships, as illustrated in Table 1 [1,4,5] with problems mainly linked with environmental protection: maritime transport accounts for about 3% of global petroleum consumption but contributes 14% of NOx and 16% of SOx.

* Corresponding author. Tel.: þ39 011 0904608; fax: þ39 011 0904699. E-mail address: [email protected] (S. Specchia). 0360-3199/$ – see front matter ª 2008 International Association for Hydrogen Energy. Published by Elsevier Ltd. All rights reserved. doi:10.1016/j.ijhydene.2008.11.092

2027

international journal of hydrogen energy 34 (2009) 2026–2042

mCO mfuel MW O2CR SCR DHv hconvH2 hconvTOT

Nomenclature cp cv HCR HHVFuel k LHVH2 LHVCO LHVFuel mH2

specific heat capacity at constant pressure, kJ kg1 K1 specific heat capacity at constant volume, kJ kg1 K1 hydrogen (H) to carbon ratio in the fuel higher heating value of the fuel, MJ kg1 expansion/compression coefficient lower heating value of H2, MJ kg1 lower heating value of CO, MJ kg1 lower heating value of the fuel, MJ kg1 H2 mass flow rate, kg h1

hOVGross hOVNet l

Other relevant pollutant emissions are particulate matter (PM), VOCs and PCAs. However, CO2 emissions are relatively low since the thermal efficiency of the engines (or combined cycle turbines) used in large vessels tend to be amongst the highest of all prime movers and of static combined cycle generating plant [6]. Under the spur of national and international committees [7–9], the main naval companies are considering the use of integrated electric plants in future naval ships. The implementation drivers are primarily lower life cycle costs, reduced noxious emissions (in particular PM and NOx), low vibrations and noise levels. The fuel cells (FCs) offer several advantages over diesels [10–12]: these include higher thermal efficiency; a flat efficiency curve vs load; lower emissions, vibrations and noise. Presently, however, FCs initial costs are significantly higher than for diesels. Hydrogen may offer considerable potential as a marine fuel. Its reduced mass when compared with existing hydrocarbon fuels (the total fuel thermal power being the same) can usefully increase the ship-owner payload; this in turn benefits the economics of oceanic transport and provides the opportunity to compete in new markets [13]. Moreover, the potential to virtually eliminate pollution at the point of use may boost significantly the

CO mass flow rate, kg h1 fuel mass flow rate, kg h1 fuel molar weight, kg kmol1 oxygen (O2) to carbon ratio steam to carbon ratio fuel vaporisation enthalpy, kJ kg1 conversion to H2 efficiency of the fuel processor conversion to H2 and CO efficiency of the fuel processor gross electrical efficiency of the overall system net electrical efficiency of the overall system O2/O2 stoichiometric in fuel burners

naval industry considering that exhaust emissions from shipping are becoming a matter of global concern. But the relatively poor volumetric energy density achieved by even the most effective storage options [14,15] can be a limiting factor for ships, particularly for the largest ones. This is particularly the case with high-speed vessels where the fuel load is proportionally greater. Given the existing fuelling infrastructure for marine transports (mainly marine diesel, kerosene and natural gas) and the relatively space restrictions on-board marine vessels, combined fuel reforming and FCs technologies are an attractive option for shipboard power systems (propulsion and/or auxiliary power requirements [16– 21]), for future ‘Green Ship’ applications (see Table 1 [22]). Moreover, they also allow tight integration of multiple thermal sources and heat loads, making them an ideal candidate for combined heat and power (CHP) marine applications [23–27]. Some of the FCs benefits provided to the process industry could also apply in the marine field [1,3–5,14,19,20]. Of special interest is the FCs high efficiency, since it may translate into fuel cost savings. Moreover, FCs efficiency is relatively constant over a broad range of power settings: such

Table 1 – Main characteristics of the major propulsion systems for various maritime transports and requirements of FCsbased APU systems. Tourist crafts Low speed diesel engine Medium speed diesel engine High-speed diesel engine Simple cycle gas turbine Advanced cycle gas turbine Mechanical propulsion Electric propulsion Fuel cells FCs-based APU systems

X

Leisure crafts

Offshore support vessels

Research and survey vessels (icebreakers)

Fast ferries

X

X

X

X

X

X

X

X

X

X

X

Ferries

X

Passengers cruise vessels

X

Coastal cargo vessels

International cargo vessels

X

X

X

X X

X

X

X

X

X

X

X

X

X

X X

X X

X X X

X X X

X X

X X X

X X X

X X

X X

2028

international journal of hydrogen energy 34 (2009) 2026–2042

a characteristic suggests that FCs might be efficiently employed in ships that frequently vary power demand – e.g., towboats, ferries, offshore supply boats, or icebreakers. FCs have few moving parts, suggesting minimal work manning requirements. They are quiet, evoking possible uses on antisubmarine warfare ships, seismic vessels and high-class yachts. Furthermore, FCs are of modular design enabling flexibility in the arrangement of plant components and could lead to a more cost-effective layout of power and cargo spaces and of basic ship structure. For naval application, the typical FC stack power ranges from 250 kW up to some MW [28]. The main challenges concerning the installation, setting and operation of such FCs power plant on board of an existing ship are quite a complex target, from several technical points of view. Two main items are to be analysed and solved: the location on board where to install the APU – to be characterised by enough free space and height and by a sufficient structural strength to sustain the overload on the selected deck – and the interconnection with the on board existing plants and facilities. An example is given in Fig. 1, showing where the APU plant could be located above the existing engine room of a just in service passenger ship (Ro-Pax vessel). For naval applications, MCFCs (Molten Carbonate Fuel Cells) and PEMFCs (Proton Exchange Membranes Fuel Cells) seem to be the best candidates [16,17,19,20,28–31]. MCFCs were chosen in the perspective of using an APU system based on a marine diesel oil processor, due to the following reasons: firstly, CO can be fed to the FC as a valuable fuel, and this is a considerable advantage when syngas from diesel reforming is chosen as the FC fuel. In fact, PEMFCs tolerate only CO traces and expensive steps of water–gas shift and preferential CO oxidation have to be foreseen. Moreover, MCFCs efficiency is higher than that of PEMFCs and does not need noble metals as catalysts. Finally, the MCFCs working temperatures are optimal for carrying out the reforming inside the vessel in case of feeding light hydrocarbons or CH4, thus exploiting the amount of heat released by the cell stack itself. Instead, PEMFCs have lower working temperatures, making this operation unfeasible. Up today, lots of progress has been done for FCs marine applications, but most of the results are still not available due to research and commercial contracts; essentially, only hint are available in literature [32–44]: so far, FCs are still in a demonstration phase.

ENGINE ROOM

Fuel Cell

On the grounds of these considerations, MCFCs were selected as the most promising technology in the marine field: in the present work a 500 kWe APU was considered, and the efficiency of such system was evaluated through the modelling of the entire process. This activity was carried out via the code Matlab/Simulink. The aim of this work was to evaluate the impact of the fuel processor on the overall net electrical efficiency, on the amount of H2 fed to the FC and on the process complexity. Two technical options were considered and compared: i) an ‘‘ATR’’ (Auto-Thermal Reforming) based system, which was also investigated by other authors [45]; ii) a ‘‘Cracking þ SR’’ (Steam Reforming) based system (cracking of the liquid fuel feed to produce light gaseous hydrocarbons followed by steam reforming into the MCFC vessel, as mentioned before). In addition, solutions to maximize the heat integration of the streams within the system were investigated, in order to increase the overall efficiency. Finally, the APU was converted into a Combined Heat and Power (CHP) system by exploiting the sensible heat of the flue gases to partially cope with the energy demand of the ship auxiliary heat utilities.

2.

ATR-based system

2.1.

Process description

The APU process taken as the reference case is the basic one described by Specchia et al. [46]. Two main parts of the whole process can be detected: the FPM and the FCM. The former is the Fuel Processor Module, that converts the diesel fuel feedstock into syngas; the latter is the Fuel Cell Module, which mainly converts the chemical power of the syngas into the electrical one. The two parts are almost independent, except for the thermal integration of the outlet turbine flue gases, which provide their sensible heat to produce steam conveyed to the reformer. The key equipment of the FPM is the AutoThermal Reformer (ATR): it is fed with desulphurized diesel, steam and pre-heated air. The pre-heating section is

FUEL OIL

Piping Connections

Fuel Processor Unit

Fig. 1 – Scheme of a possible allocation of the FC-APU units on-board large Ro-Pax ships (courtesy of Cetena S.p.A.).

2029

international journal of hydrogen energy 34 (2009) 2026–2042

a Ex-03

Ex-04

S-101 Feed-01

Fuel-05

Demi-water-03

Steam/demi-water tank

Air-05

FCS

S-102

K-102

Air-06

APS

Air-04

b

Ex-01

Cath-04

An-01

FB-101

Cath-07

Ref-04

Ref-03

Anode

AIR

CB-101

Air-07

DEMI WATER

F-101

Cath-06

E-108

Demi-water-01

P-102

An-02

Cathode

Cath-05

E-107

Cath-02

Ref-05

Ref-02

Cath-01 An-03

An-04

Cath-03

E-106

Fuel-03

Steam-04

E-102

Demi-water-04

Ref-01

ATR

Air-01

AIR

EXHAUST GASES

Steam-01

V-101 Demi-water-02

Air-02

K-101

Exhaust-06

X-101 Feed-02

X-102

EXHAUST GASES

Fuel-01

DIESEL

P-101

Ex-05

Fuel-02

Steam-02

Steam-03

Air-03

Fuel-06

Air-08

Fuel-04

Ex-02

P-88 AIR

E-105

E-103

E-101

FB-102

F-102

T-101

139.8 kg h-1 (LD) @ 25°C; 0 barg

4179 kg h-1 (air) @ 25°C; 0 barg

-1

371 kg h (demi-W) @ 25°C; 0 barg

541 kWe

OVnet =

32.7%

1113 kW wasted

+ APS

Light diesel JP-5

879 kWe 746

?OV 39.0% OVgross gross= 45.1%

APU

93 kg h (SR) 26.3 kg h-1 (FB-101) -1 20.5 kg h (FB-102)

POWER

-1

HEAT

1654 kW

P AU + GT GTAPU APU + -APS

224 kW kWe 205

-1

3826 kg h @ 447°C (GTAPS exhaust gases) + 864 kg h-1@ 209°C (FB-102 flue gases) + heat losses

864 kg h-1 (4.6%mol H2O) (FB-102 flue gases) @ 209°C; 5 mbarg

-1 3826 kg h (17.8%mol H2O) (GTAPS exhaust gas) @ 447°C; 5 mbarg

Fig. 2 – (a) Scheme of the reference ‘‘ATR’’ based system and (b) mass/energy balances.

constituted by a flame burner which pre-heats steam and air up to the inlet temperature of the ATR, and pre-heats as well diesel to operate a liquid desulphurization. The FCM must be provided with an extremely low sulphur content stream (around 0.1 ppm), and this can be accomplished through

a second step of desulphurization, this time in gas phase. The gas is then cooled down for the desulphurization and reheated again before to enter ATR as displayed in the schematic of the process (Fig. 1 in [46]). In the FCM the electrical power is generated by two routes:

2030

international journal of hydrogen energy 34 (2009) 2026–2042

i) in the FCS (Fuel Cell System) the syngas is electrochemically converted into a MCFC with a nominal power of 500 kWe; ii) in the APS (Air Process System) the outflow of the MCFC is expanded into a gas turbine, thus generating approximately one fifth of the overall FCM electrical power. In the present work, the FCS described in [46] was redesigned (streams nomenclature is consistent with Fig. 2a showing the overall process): Ansaldo Fuel Cells [47] manufactured a new FCS with four MCFC stacks, each one composed by 150 cells in parallel and providing a nominal power of 125 kWe. The stacks (ST-01, ST-02, ST-03 and ST-04) are inside a pressurized vessel at 2.6 barg (identified by the dashed line of the FCS in Fig. 2a) and are fed in parallel with the H2–CO rich reformate gas (Ref-05). A detailed scheme of the pressurized vessel is depicted in Fig. 3a. The unreacted fuel out-flowing from the stacks’ anodes (An-01) is conveyed to a catalytic burner (CB-101) to complete the combustion. The outlet stream from CB-101, rich of CO2, is sent to the stacks’ cathode sides, through the blower F-101 together with the air rate Air-07 coming from the compressor K-102; the temperature control of the cathode sides feed is obtained through the heat exchanger E-108 with steam production. The outlet streams from the cathode sides Cath-01 go directly into the pressurized vessel, surrounding all the equipments here installed: MCFC stacks, catalytic burner, fan, heat exchanger. The gaseous atmosphere inside the pressurized vessel is partly fed to CB-101 through F-101 (the stream rate Cath-02 is about 80% of the total cathode side outlets Cath-01). The remaining part of the cathodic atmosphere (Cath-03) goes directly outside the pressurized vessel, to the APS system. This configuration ensures a high CO2 concentration into the cathode side feed. However, compared to the FCS presented in [46], the recirculation entails a higher energy consumption for the blower F-101 and larger volumes of the involved equipments. As far as the APS is concerned, the gas turbine (T-101 in Fig. 2a) expands the flue gases from the FCS (Cath-03), after the temperature increase in the direct combustor FB-101. The compressor K-102, driven by the turbine, provides air for FCS. Part of the compressed air is conveyed to the MCFC (Air-07), while the remaining part (Air-06) is mixed with Cath-05 and processed into the direct burner FB-101 to increase the temperature; the resulting gas mixture is then expanded into the gas turbine. As can be observed in Fig. 2a, the stream Cath04 can be used as a bypass of the turbines: the flow rate of the bypassed stream is tailored for an acceptable balance of the APS rates, i.e., the rate expanded in the turbine with the one compressed in K-102 (maximum Cath-06/Air-04 mass ratio equal to 1.2). In order to ensure the best oxygen to fuel ratio in the burner FB-101 despite the variation of the Cath-05 flow rate, Air-06 is regulated accordingly. The overall efficiency of the process at different operating conditions was evaluated using the following equations and assumptions for the equipment modelling.

2.2.

Models description

2.2.1.

Feedstocks

The raw materials of the process were light diesel, air and demineralised water. All the reactants were considered to be

fed at 25  C and 0 barg. JP5 grade was selected as a suitable light diesel for commercial ships, and its properties of interest for the simulation are listed in Table 2 [48]. The selection of this standard as the fuel for the process accounted for the balance between its cost and the one associated to equipment for the desulphurization. As far as water is concerned, it could be both fed by an external make-up and recycled through the condensation of the turbine outlet flue gases, characterised by a huge steam concentration. This solution is described in [46]. No timeframe dynamics were investigated, thus mass and energy balances were implemented in steady-state.

2.2.2.

Liquid desulphurization

Diesel liquid desulphurization (S-101 in Fig. 2a) was performed through reactive adsorption of the sulphuric compounds, at around 200  C via Ni-based sorbents (the principle of this operation is described for example in [49,50]), aiming at reducing the sulphur content to acceptable levels for the reforming catalyst (i.e. around 50 ppm). In the present case, the operation was carried out under a pressure of 9 barg via the liquid diesel P-101 and at a constant temperature of 185  C (fuel was pre-heated in E-105).

2.2.3.

Flame burner

The pre-heating section of the process was made of a cascade of heat exchangers whose heating fluids were the flue gases flowing out from the flame burner (FB-102 in Fig. 2a): a certain amount of fuel (Fuel-06) was split from the main flow and conveyed to this equipment, where combustion with air (Air-08) took place. This burner had an output temperature around 960  C: this temperature decreased as long as air, steam and liquid fuel were heated up. These three streams were mixed (X-101 and X-102) and sent to the ATR at around 603  C. Fundamentals of such device can be found in [51].

2.2.4.

Auto-thermal reformer

In this flow scheme the fuel processor was constituted by an auto-thermal reformer, whose application as a means of producing syngas for fuel cell exploitation is widely described in literature, such as [24–27,46,52]. The reactions involved were: Combustion : C11:96 H22:21 þ 17:51O2 /11:96CO2 þ 11:11H2 O

(1)

Reforming : C11:96 H22:21 þ 11:96H2 O411:96CO þ 23:07H2

(2)

Water–Gas Shift : CO þ H2 O4CO2 þ H2

(3)

Methanation Equilibrium : CO þ 3H2 4H2 O þ CH4

(4)

The outlet composition was calculated at the thermodynamic equilibrium, through the Gibbs free energy minimization method (described in [53]). At temperatures compatible with this process, the methanation equilibrium gave little effect. The selected steam to carbon ratio (SCR) was 3.1 in order to avoid soot formation from the fed fuel cracking. This choice differs from the value suggested by Specchia et al. [46], since it was evaluated to be more energy effective (a lower water mass flow rate has to be pre-heated, thus saving fuel in the flame burner), but it is still enough towards H2 production

international journal of hydrogen energy 34 (2009) 2026–2042

a

Anode-01 Cath-01.1 Cathode-01

Cath-02 An-01

Ref-05

from T ATR

Anode-02 Cath-01.2 Cathode-02

Stack-01

Stack-02

An-02

CB-101

Stack-03

Stack-04

Anode-03

Anode-04

Cathode-03

Cathode-04 Cath-01.3

Cath-01.4

Air-07

F-101

An-03

Demi-water-05

E-108

Steam-04

Tfrom/to Steam/Demiwater tank

An-04

Cath-03

MCFC VESSEL

Tto FB-101

Tfrom K-102

b Tfrom Cracking

Anode-01 Cath-01.1 Cathode-01

Cath-02 An-01

Crack-04

Anode-02 Cath-01.2 Cathode-02

Stack-01

Stack-02

CB-101

An-02

SR-101

Stack-03

Stack-04

Ref-01

Anode-03

Anode-04

Cathode-03

Cathode-04 F-101

An-03

Cath-01.4

Air-04

Cath-01.3

Demi-water-05

E-108

Steam-04

Tfrom/to Steam/Demiwater tank

An-04

MCFC VESSEL

Cath-03 Tto FB-101

Tfrom K-102

Fig. 3 – Scheme of the pressurized vessel containing the MCFC stacks: (a) with ‘‘ATR’’; (b) with ‘‘Craking D SR’’.

2031

2032

international journal of hydrogen energy 34 (2009) 2026–2042

Table 2 – JP5 grade physical and chemical properties. Fuel

Formula

HCR MW DHv [kJ kg1] Cpl [kJ kg1 K1] Cpv [kJ kg1 K1] HHV [MJ kg1] LHV [MJ kg1] Sulphur [wt%]

JP5

C11.96H22.21

1.86

166

270

2.2

1.7

maximisation and safe regarding to coke formation. The fuel processor conversion efficiency was calculated as follows: hconvH2 ¼

_ H2 ,LHVH2 m _ FUEL ,LHVFUEL m

hconvTOT ¼

_ H2 ,LHVH2 þ m _ CO ,LHVCO m _ FUEL ,LHVFUEL m

(5)

(6)

The outlet ATR temperature was around 776  C. The flow was then cooled down to 320  C (in order to perform the gas desulphurization in S-102), by pre-heating the inlet demiwater (E-107). Finally, the H2 and CO rich stream was fed to the FCM (Ref-05). It was not necessary to eliminate CO, as in the case of PEMFCs [52], being a valuable fuel for MCFCs. Therefore the operating conditions were optimized such as to maximize hconvTOT.

2.2.5.

Gas desulphurization

Gas desulphurization (S-102) was carried out at temperatures well below the ATR outlet and the FCM ones; therefore an undesired cooling down and then heating up of the gas stream was needed (E-106). The reformate was treated at around 320  C via the established technology of ZnO sorbents to eliminate H2S (below 0.1 ppmv) [54]. Despite their high uptake capacity, these sorbents suffer from two main drawbacks: thermal stability and competitive adsorption between H2O and H2S. Hence, the presence of steam influences negatively the sorption equilibrium of H2S. This step was considerably modified by the ‘‘Cracking þ SR’’ based system, thus improving the overall process efficiency.

2.2.6.

Fuel cell system

The inlet temperature to the vessel where the stacks were located was 625  C (Ref-05), while the MCFC was maintained at 650  C by regulating the recycle into the vessel of the cathode outlet stream and the mixture with the fresh cathode inlet stream. The electrochemical reactions at the anode and cathode sides were: Anodic oxidation:  H2 þ CO2 3 /H2 O þ CO2 þ 2e  CO þ CO2 /2CO þ 2e 2 3

(7)

44.8

(8)

The anode utilization factor, namely the ratio between the total and the converted H2 þ CO into the MCFC stacks, was taken as 1.333. The remaining syngas species were burnt in a catalytic combustor (CB-101), to which the cathode recycle was conveyed as well. The CO2 supply to the cathode side was performed by recycling the outlet stream of the catalytic burner to the stacks, after mixing it with O2 rich fresh air. The pressure inside the cells was 2.6 barg and no thermal dissipation out of the vessel was taken into account.

<0.3

As far as the cell efficiency is concerned, the total 500 kWe were obtained via 4 stacks of 125 kWe, each one made of 150 cells; the single cells worked at tension of 775 mV and a current density of 1550 mA m2, resulting in a power density of 1.17 kW m2 and a net electrical efficiency of 63%. This value was representative of MCFCs working at full load in steady-state, and it was used to simulate the overall process efficiency. The working parameters arose from the data sheet referred to a plant manufactured by Ansaldo Fuel Cells [47].

2.2.7.

Air process system

The air process system was conceived as a section to exploit the enthalpy of the stream out-flowing from the FCS vessel. Hence, a turbine (T-101) expanded these gases from an inlet temperature of 860  C reached by means of the auxiliary burner (FB101). The turbine drove a compressor (K-102), which provided air at 2.6 barg to the FCS vessel. Depending on the operating conditions, around 100–120 kWe were obtained from the APS, considering that the flow to the turbine was assumed to exceed no more than 20% of the one to the compressor. The modelling of the compressor and the turbine was carried out by considering a polytropic compression/expansion with an exponent k ¼ cp/cv according to the cp and cv values of the stream at the inlet and outlet temperatures. The efficiency values of the devices are listed in Table 3.

2.2.8.

Auxiliary equipment

The other devices not mentioned so far were: the pumps, the heat exchangers and the steam generator (E-102 in Fig. 2a). Their operating parameters as well as the ones of the former devices are listed in Table 3. A minimum DT of 15, 20 and 40  C was used for liquid–liquid, gas–liquid and gas–gas heat exchangers, respectively.

2.2.9.

Process efficiency

The total net electrical power accounted for the power produced by the FC and the turbine (Gross Power), minus all the electrical parasitic powers (Compressors and Pumps). The ratio between this quantity and the chemical power of the fuel processed to the system gives the overall net electrical efficiency hOVNet. The expressions of the Gross and Net electrical efficiencies were calculated as follows:

Cathodic reduction: O2 þ 2CO2 þ 4e 42CO2 3

42.5

hOVGross ¼

Gross Power _ FUEL ,LHVFUEL m

hOVNet ¼

Net Power _ FUEL ,LHVFUEL m

2.3.

Fully integrated system

(9)

(10)

The overall net electrical efficiency of the system can be considerably improved by reducing the value of the denominator in Eq. (10). This could be achieved through a higher thermal integration of the outlet streams and the pre-heating

international journal of hydrogen energy 34 (2009) 2026–2042

Table 3 – Parameters assumptions used in the model calculations. Fuel Pump P-101 Inlet/Outlet pressure [barg] Efficiency Water Pump P-102 Inlet/Outlet pressure [barg] Efficiency ATR Air Compressor K-101 Inlet/Outlet pressure [barg] Efficiency Cooled outlet to 80  C Stack Air Compressor K-102 Inlet/Outlet pressure [barg] Efficiency Stack Turbine T-101 Inlet/Outlet pressure [barg] Efficiency Burner Blower F-102 Inlet/Outlet pressure [barg] Efficiency Flame Burner FB-102 Lambda Outlet temperature [ C] Liquid Desulphurizer S-101 Operating Pressure [barg] Operating Temperature [ C] Gas Desulphurizer S-102 Operating Pressure [barg] Operating Temperature [ C] ATR System O2CR SCR Pressure [barg] Outlet composition at the thermo-dynamic equilibrium Vapour Producer V-101 Steam production at 449  C and 8.3 barg Heat losses ¼ 10% total heat duty MCFC Anode Utilization Factor Cell Temperature [ C] Cell Pressure [barg] Cell Efficiency Cathode Recycle MCFC Blower F-101 Inlet/Outlet pressure [barg] Efficiency Auxiliary Burner FB-101 Lambda Outlet temperature [ C] Heat Exchangers Heat losses ¼ 10% heat exchanger duty

0/9 0.7 0/8.3 0.7

system and a possible way to overcome this problem is to use the flame burner as an auxiliary or a spare equipment, with a quite fast response in case of need.

3. Cracking unit D steam reformer based system 3.1.

0/7 0.65

0/2.6 0.8 2.6/0.1 0.8 0/0.5 0.7 2.84 960 9 185 2.55 302 0.35 3.1 2.6

1.333 650 2.55 0.63 0.8 2.6 0.9 4.7 860

section of the ATR feed. A schematic of the proposed improvements is depicted in Fig. 4a. The economy arising from such integration allowed removing the flame burner and saving the associated fuel. In addition, the burner blower (F-102 in Fig. 2a) was no longer needed. The outlet turbine flue gases were used to pre-heat air (E-101), then water was heated up to saturation (E-104), vaporized (E-103) and after that super-heated (E-102) in a cascade of heat exchangers, and finally fuel was heated to the liquid desulphurization temperature (E-105). The partial drawback of such integration was a reduced flexibility of the

2033

Process description

The main concept behind this system (depicted in Fig. 5a) is to split the syngas production into two steps: firstly, a cracking of the already desulphurized liquid diesel to obtain a mixture of light gas compounds, with a number of C atoms between 1 and 6, and a heavy liquid as a residue. The operation should be at a moderate temperature, such as 350–400  C, and not too energy consuming. If the step is efficient enough, and a good conversion to gas species is obtained, a steam reforming at higher temperature is carried out to produce H2 and CO. This was done by integrating the steam reactor (SR-101) inside the pressurized stack vessel (see Fig. 3b), where temperatures are compatible with this operation: hence, SR-101 was thermally coupled to the catalytic burner (CB-101) to undergo syngas production from the gaseous output of the cracking unit (Crack-04). As a result, the heat produced inside the fuel cell, instead of being rejected for the major part through E-108 as in the ‘‘ATR’’ based scheme, is exploited to drive the steam reforming reaction. In addition, gas desulphurization (S-102) could be placed between the two steps, at a suitable temperature (320  C), very close to the cracking outlet temperature (360  C), as depicted in Fig. 5a. This approach solved some limitations of the ATR-based system very efficiently: the heat exchanger E-106 in the ATR process scheme (Fig. 4a) could be completely removed. In addition, water for steam reforming was placed after the gas desulphurization, avoiding water inhibition effect in the H2S adsorption equilibrium. In this case it is possible to employ inexpensive adsorbents of the ZnO family as potential candidates to get down to around 0.1 ppmv before the reformer. Additional advantage could be the possibility to thrift the load of noble metals in the reforming catalysts, as a consequence of the reduced sulphur content and the reduced coking tendency of the feed gases. From the efficiency point of view, this configuration was particularly effective: the heat produced by the conversion of H2 and CO, and the one released in the catalytic burner, were exploited to carry out the reforming reaction. As far as heat integration of the FPM and the FCM was concerned, the enthalpy of the outlet turbine flue gases was used to pre-heat the fuel feed of the cracking unit and to vaporize the water necessary for the steam reforming up to 350  C. Thus, the heat duty involved in this step was not too large, far below the one of the case described in Section 2.3.

3.2.

Models description

3.2.1.

Cracking unit

The cracking unit was modelled by taking as a reference example the work of [55]. The main components of the cracking gas output are: H2, CH4, C2H4, C2H6, C3H6, C3H8, C5s, BTX. The

2034

international journal of hydrogen energy 34 (2009) 2026–2042

a

E-101

E-102

E-103

E-105

E-104

Ex-03

Ex-04

Ex-05

Ex-06

EXHAUST GASES

Ex-01

V-101

Fuel-03

CB-101

Air-07

S-102

K-102 Air-04

b

Air-06

Air-05

FCS

AIR

Cath-04

An-01

APS

FB-101

Cath-07

Anode

Cath-05

Ref-04

P-102

Ref-03

DEMI WATER

F-101

Cath-03

Ref-05

Ref-02 Demi-water-01

An-02

Cathode E-108

Cath-02

Cath-01 An-03

An-04

Cath-06

Steam/demi-water tank Demi-water-05

ATR

Steam-04

Steam-02

E-106

E-107

Fuel-04

X-101

Demi-water-02

Feed-02

X-102

Ref-01

Air-01

AIR

S-101

Fuel-05

Feed-01

K-101

Fuel-01

Demi-water-04

Air-02

P-101

Steam-01

Steam-03

Air-03

DIESEL

Demi-water-03

Fuel-02

Ex-02

T-101

-1

94 kg h (SR) 26.3 kg h-1 (FB-101) Light diesel JP-5

POWER

879 kWe 749

?OV 39.0% OVgross gross= 52.6%

HEAT

1423 kW

APU+ APU GT +T-APS

224 kW kWe 192

655 kWe 557

?OV 29.0% net= 39.1% OVnet

866 kW wasted 3826 kg/h @ 290°C (APS exhaust gases) + heat losses

3335 kg h-1 (air) @ 25°C; 0 barg

371 kg h-1 (demi-W) @ 25°C; 0 barg

APU + APS

-1

120.3 kg h (LD) @ 25°C; 0 barg

-1 3826 kg h (17.8%mol H2O) (APS exhaust gas) @ 290°C; 5 mbarg

Fig. 4 – (a) Scheme of the thermally integrated ‘‘ATR’’ based system and (b) mass/energy balances.

working conditions were chosen in the range of an upper conversion limit of 0.8 (defined as the ratio between the outlet mass of gaseous compounds and the total inlet mass) at 360  C, and a lower conversion limit of 0.6 at 340  C. The heat for the cracking operation was given by a small flame burner (FB-102), designed with the duty to also increase the temperature of the steam reforming feed from 350  C up to 677  C. The anode was fed at 625  C, as in the ATR-based system. The liquid residue was used as a valuable fuel for the auxiliary burner in the APS, thus reducing the amount of fresh fuel directed to this burner.

3.2.2.

Steam reformer

The reactions involved in the modelling of the steam reforming (SR-101) of gaseous species from cracking of light diesel were: i) Reforming reactions, with x and y varying according to the cracking gas products:
(11) Cx Hy þ xH2 O4xCO þ ðx þ y=2 H2 ii) Water–gas shift and methanation equilibrium according to equations (3) and (4). In this case, the methanation

2035

international journal of hydrogen energy 34 (2009) 2026–2042

a

E-101

E-102

E-104

E-103 Ex-03

Ex-04

Ex-05

EXHAUST GASES

Fuel-02

Ex-02

Fuel-01 Demi-water-04

Fuel-04

S-101

EXHAUST GASES

Ex-08

Fuel-03

Steam/demi-water tank

Ex-01

Steam-04

V-101

Steam-02

Carck-01

DEMI WATER

S-102 Demi-water-05 Ex-07

Air-02

E-106

K-102 Air-01

b

Cath-04

An-01

FCS

AIR

CB-101

Cath-05

Anode

F-101

Air-03

APS

FB-101

Cath-07

Ref-01

E-108

Cath-03

An-02

Cathode

Cath-02

An-03

Air-04

X-101

SR-101 An-04

Crack-04

Crack-03

Steam-03

Cath-01 Crack-02

Ex-06

P-102

Cath-06

Fuel-07

Fuel-05 Steam-01

FB-102

Air-05

Demi-water-03

F-102

Cracking

Fuel-06

P-101

AIR

Demi-water-02

DIESEL

T-101

-1

112 kg h (LD) @ 25°C; 0 barg

-1

3507 kg h (air) @ 25°C; 0 barg

-1

345 kg h (demi-W) @ 25°C; 0 barg

HEAT

POWER

879 kWe 756

?OV 39.0% OVgross gross= 57%

APS

-1

APU +

1325 kW 89.7 kgkg/h h (ATR) 121.9 (ATR) 22.3 kg/h kg h-1(FB (FB-101) 23.8 45.2 kg/h (FB of Light Diesel Light diesel (2255 kW)JP-5

APU APU + GT GT+ -APS

165 kW kWe e

655 kWe 591

?OV 44.6% OVnet net= 29.0%

734 kW wasted 507 kg h-1 @ 374°C (FB-102 flue gases) + -1 3457 kg h @ 373°C (APS exhaust gases) + heat losses

-1

507 kg h (4.6%mol H2O) (FB-102 flue gases) @ 374°C; 5 mbarg

-1 3457 kg h (16.3%mol H2O) (APS exhaust gas) @ 373°C; 5 mbarg

Fig. 5 – (a) Scheme of the reference ‘‘Cracking D SR’’ based system and (b) mass/energy balances.

equilibrium was shifted to the left hand side as a veritable reforming, since methane was present in the inlet flow. A SCR equal to 4 was used to drive the steam reforming process.

3.3.

Fully integrated system

A highly integrated system could be obtained, aiming to maximize the overall net electrical efficiency of the process. This could be achieved as arises by the schematic of Fig. 6a.

The turbine flue gases, instead of being involved only in the fuel and steam pre-heating section, were closely integrated as heat source in the anode pre-heater. However, due to temperature driving force limitations in E-106 (the inlet temperature of the heating stream was in this case lower than that of Ex-06 in Fig. 5a, deriving from diesel oil flue gases), they were not able to let the steam reforming feed reach 677  C, as in the former scheme of Fig. 5a, but only 625  C; the highest reachable value was, therefore, somewhat 50  C lower (see Table 5). The residual heat to enter into the stack at the same

2036

international journal of hydrogen energy 34 (2009) 2026–2042

E-103

Demi-water-02

S-101

Ex-03

Fuel-06

Fuel-04

Fuel-01 DIESEL

E-102 Ex-02

P-101

EXHAUST GASES

Ex-04

Demi-water-04

Fuel-02

Ex-08

Steam-01

E-101

E-104

Demi-water-03

a

Steam-02

S-102 Ex-07

E-108 Anode

An-01

F-101

Cath-03

An-02

Cathode Air-04

Crack-04

An-03

CB-101

X-101 E-107

E-106

K-102

Air-01

AIR

Air-03

APS

FB-101 Cath-06

Cath-05

Cath-04

FCS

Air-02

Crack-03

SR-101 An-04

Ref-01

Ex-06

Cath-01

Crack-02

Ex-05

P-102

Cath-02

DEMI WATER

Fuel-03

Carck-01

Demi-water-05

Steam/demi-water tank

Steam-04

V-101

Fuel-07

Cracking

Steam-03

Fuel-05

T-101 Cath-07

Ex-01

b

3012 kg h-1 (air) @ 25°C; 0 barg

-1

345 kg h (demi-W) @ 25°C; 0 barg

+ APS

100 kg h-1 (LD) @ 25°C; 0 barg

655 598 kW kWe

e ?OVnet = 29.0% OV net= 50.6%

585 kW wasted -1

3457 kg h @ 283°C (APS exhaust gases) + heat losses

3457 kg h-1 (16.2%mol H2O) (APS exhaust gas) @ 283°C; 5 mbarg

APU

Light diesel JP-5

POWER POWER

89.7 kg h-1 (SR) -1 10.3 kg h (FB-101)

879 757 kW kWe

e ?OVgross = 39.0% OV gross= 64%

HEAT HEAT

1183 kW

GTAPU APU + GT + -APS

224 kW e kWe 159

Fig. 6 – (a) Scheme of the reference thermally integrated ‘‘Cracking D SR’’ based system and (b) mass/energy balances. temperature as in the former case (650  C) was obtained by a higher heat duty supplied to SR-101 from the catalytic burner CB-101 (see in Table 5 the SR-101 heat duty increase to 130 kW as compared to 115 kW of the former case). Finally the heat make-up to restore the same inlet temperature to the cathode inlet was obtained by heating the fresh compressed air conveyed to the FCS (Air-02). This operation was fulfilled by the turbine flue gases in E-107, after the first service in E-106 downstream the cracking unit. Finally, the turbine flue gases were used to pre-heat the fuel and steam feeds.

4.

Discussion

4.1.

Simulation results

The simulation results are listed in Table 4, where the ‘‘ATR’’ based system and the ‘‘Cracking þ SR’’ based one are compared. Table 4 displays as well the simulation results for the two fully integrated schemes. These data were used to calculate the gross and net electrical efficiencies of the system. In addition, Table 5 shows the operating conditions

2037

international journal of hydrogen energy 34 (2009) 2026–2042

Table 4 – Simulation results for the ‘‘ATR’’ and the ‘‘Cracking D SR’’ based systems.

Fuel Pump P-101 Required power [kW] Water Pump P-102 Required power [kW] ATR Air Compressor K-101 Required power [kW] Stack Air Compressor K-102 Required power [kW] Stack Turbine T-101 Obtained power [kW] Burner Blower F-102 Required power [kW] Vapour Producer V-101 Heat duty [kW] MCFC Electricity Duty [kW] Heat Duty [kW] MCFC Blower F-101 Required power [kW] Auxiliary Burner FB-101 l (O2/O2stoich) Outlet temperature [ C] Streams [kg h1] Fuel to FPM Fuel to Flame Burner FB-102 Fuel to Auxiliary Burner FB-101 Water to FPM Air to FPM Air to APS Compressor K-102 Air to Flame Burner FB-102

‘‘ATR’’-based scheme

Fully integrated ‘‘ATR’’-based scheme

‘‘Cracking þ SR’’ based scheme

Fully integrated ‘‘Cracking þ SR’’ based scheme

0.063

0.054

0.014

0.013

0.123

0.123

0.035

0.035

33.5

33.5

137.5

137.5

137.5

137.5

249.4

249.4

254.3

255.6

12.2

7.2

239.1

272

207

207

500 158

500 160

500 3.6

500 4.9

21.1

21.1

20.9

20.9

4.7 860

4.7 860

4.8 882

4.8 887

93 20.5 26.3

94 0 26.3

89.7 12 22.3

89.7 0 10.3

371 323.4 3012 843.8

371 323.4 3012 0

345 0 3012 495

345 0 3012 0

and stream compositions for the ATR, the cracking unit and the SR, both in the basic and fully integrated schemes. The ‘‘ATR’’ based system (depicted in Fig. 2b) provided 746 kWe (500 kWe from the MCFC and 246 kWe from the gas turbine T-101), of which only 541 kWe were available, since the remaining power accounted for the balance of plant requirements: these were firstly the stack air compressor K-102 (137.5 kW), the MCFC blower F-101 (21.1 kW), the ATR air compressor K-101 (33.5 kW) and the burner blower F-102 (12.2 kW), while the other utilities’ requirements were almost negligible (P-101, P-102). As a result, the system had a gross electrical efficiency equal to 45.1% and a net electrical efficiency equal to 32.7%. The thermal integration of the system was limited to steam production in evaporator E-102. The chemical energy ‘‘stored’’ in the fuel was wasted by several means: as sensible heat in the outlet flue gases (Ex-05 and Ex06), as heat losses in the heat exchangers (10% of the total heat duty exchanged), as heat losses inside the equipment (requiring energy for compression) and as rejected heat of cooling systems (E-108 for temperature regulation inside the FCS pressurized vessel). In particular, the last contribution was quite severe for this system (158 kW), and reduced considerably the effectiveness of this configuration. Such result pushed us to investigate the solution with the inside vessel SR. The integration of the exhaust flue gases with the preheating section of the system arose from the possibility of

reducing the amount of wasted kWth in the outlet streams (see Fig. 4b), thus increasing the net electrical efficiency up to 39.1%. Hence, the amount of heat associated to the flame burner FB-102 was coped by the outlet turbine flue gases (Ex-01 in Fig. 4a): the elimination of F-102 (12.2 kWe) increased the net electrical efficiency while the thermal integration saved 20.5 kg h1 of Fuel-06 (242 kWth). The inlet temperature to the ATR was slightly lower, but this had almost no impact on the reachable hydrogen concentration at the ATR outlet (Table 5). However, ATR was very fuel consuming since an O2/ C ratio of 0.35 was necessary to foster the steam reforming reaction. Temperatures inside the ATR were the highest ones in the system (see Tables 3 and 5), so no thermal integration was possible. For this reason, the solution to this issue was to exploit the rejected heat by FCS to drive the reforming reaction, as described in the ‘‘Cracking þ SR’’ system. What appears outstanding is that the choice to perform the reforming inside the MCFC pressurized vessel made the overall system more efficient, having a net electrical efficiency of 44.6%. The integration of the steam reformer reduced the amount of heat via the heat exchanger E-108 from 158 kW (‘‘ATR’’ based system) down to 3.6 kW (see Table 4). Moreover, the feed pre-heating was less demanding as compared to the ‘‘ATR’’ based scheme, due to the mild temperatures of the diesel cracking. The fuel economy was considerable: the diesel mass flow rate decreased from 139.8 kg h1 in the ‘‘ATR’’ based scheme (Fig. 2b) to 112 kg h1 (Fig. 5b), representing a saving of

2038

international journal of hydrogen energy 34 (2009) 2026–2042

Table 5 – Comparison between the ATR, Cracking Unit and SR outlet gas compositions.

Inlet–Outlet temperature [ C] Outlet composition [%] H2 H2O CO CO2 CH4 N2

Inlet–Outlet temperature [ C] Heat Duty [kW] Outlet composition [%] CH4 H2 C2H4 C3H6 C 4s C2H6 C3H8 C 5s BTX

Inlet–Outlet temperature [ C] Heat Duty [kW] Outlet composition [%] H2 H2O CO CO2 CH4

ATR

ATR in fully integrated scheme

603–776

500–701

28.6 34.6 6.4 9.4 0.01 21.0

29.3 33.7 5.5 10.3 0.3 20.9

Cracking unit

Cracking unit in fully integrated scheme

180–360 18.5

180–360 18.5

22.4 9.7 24.3 16.1 7.8 5.6 1.0 12.6 0.5

22.4 9.7 24.3 16.1 7.8 5.6 1.0 12.6 0.5

a full thermal integration, i.e., with a maximized net electrical efficiency) were able to provide, respectively, 10,008 kg h1 and 8328 kg h1 of hot sanitary water, thus reaching CHP efficiencies of 74.1% and 85.6%, respectively (see Fig. 7). In the CHP configuration, there is also the possibility of recycling some of the condensed water when the turbine flue gases were cooled down to the dew point (in our case fixed at 39  C). However, contrary to the original scheme (Fig. 3 in [46]), a water self-sustaining system was not possible and an amount of demi-water was in any case fed to the system. This was due to the lower water content in the turbine flue gases of our systems. Still, the external make-up was reduced to 77.3 kg h1 and 97 kg h1, respectively, representing only 20.8% and 28.1% of the original water needs. Fig. 8 depicts how the two thermally integrated systems (ATR in Fig. 8a and ‘‘Cracking þ SR’’ in Fig. 8b) were modified to take into the account the demi-water recycling and the sanitary water production in the CHP configuration.

4.2.

Steam reforming

Steam reforming in fully integrated scheme

677–650 115

625–650 130

40.5 42.6 4.5 9.9 2.5

40.5 42.6 4.5 9.9 2.5

On-board compatibility of the system

Several categories of ships were investigated, in order to identify the most common on-board auxiliaries and to evaluate the consistency between the APU requirements in terms of fuel/water and the on-board ship capabilities. The selected ship types for this analysis were: i) Cruise ferries; ii) Ro–Ro ships (ferries designed to carry containers and wheeled cargo such as trucks, cars, etc.); iii) Ro-Pax ships (ferries designed for combined vehicle transport and passenger accommodation). The contributions of the electric generation system on a ship to the total power requirements were given by: deck and 100 90 80

CHP

70 60

CHP

50 40

ATR

ATR int

Cracking+SR

Cracking+SR int

ATR int

30

Fig. 2

Fig. 4

Fig. 5

Fig. 6

Fig. 8a

20 10 0

Cracking+SR int

Efficiency [%]

20%. This saving could be maximized with the same strategy as described above, i.e., the thermal integration of the flue gases with the pre-heating section (see Fig. 6a). The final electrical efficiency of the fully integrated ‘‘Cracking þ SR’’ based scheme reached 50.6%, and the fuel consumption was lowered to 100 kg h1 (Fig. 6b): this represents a fuel saving of 28% with respect to the ‘‘ATR’’ based scheme (Fig. 2b), and still an improvement of 11% compared to the corresponding scheme with a lower degree of thermal integration (Fig. 5b). The evidence of these considerations is depicted in Fig. 7, which gathers the net electrical efficiencies of the ‘‘ATR’’ and ‘‘Cracking þ SR’’ based schemes, along with their fully integrated versions. As can be verified by the energy balances in Figs. 4b and 6b, a considerable amount of heat was still wasted in the flue gases streams, which in principle could be coupled to ship facilities, e.g., to produce sanitary water as well as hot air for on-board heating necessities. The APU system could operate then as a Combined Heat and Power CHP system, as already discussed in [46]. If sanitary water at 70  C is produced with this CHP approach, then the two schemes (ATR and ‘‘Cracking þ SR’’ with

Fig. 8b

Fig. 7 – Comparison between the overall electrical (red) and CHP (green) efficiencies of the ‘‘ATR’’ and ‘‘Cracking D SR’’ based systems, in their reference (Figs. 2–4) thermally integrated (Figs. 5 and 6) and CHP (Fig. 8a and b) configurations.

2039

international journal of hydrogen energy 34 (2009) 2026–2042

a

E-101

E-102

E-103

E-105

E-104

Ex-03

Ex-04

Ex-05

E-109

Ex-06

EXHAUST GASES

Ex-07

Feed-02

Demi-water-06

Fuel-04

S-101

Fuel-05

Feed-01

SANITARY WATER

Fuel-01

Demi-water-04

10008 kg h-1 25°C

Steam-01

Air-02

P-101

Demi-water-02

Steam-03

Air-03

DIESEL

Demi-water-03

Fuel-02

Ex-02

X-102

X-101 Sanitary-water-01 Steam-02

10008 kg h-1 70°C SANITARY WATER

Sanitary-water-02

V-101

K-102 AIR

E-108 Anode

An-01

F-101

Cath-07

Sanitary water-02

Demi-water-06 Cath-03

An-02

Cathode Air-04

Ref-01

Crack-04

An-03

Cath-02

Ex-07

SR-101 An-04

CB-101

X-101 E-107

E-106

Air-01

K-102

FB-101

Air-03

APS

Ex-01

Fig. 8 – CHP schemes of the ‘‘ATR’’ (a) and ‘‘Cracking D SR’’ (b) based systems.

Cath-06

Cath-05

Cath-04

FCS

Air-02

Crack-03

Crack-02

Ex-05

Ex-06

Cath-01

Demi-water-01

S-102

P-102

AIR

V-102 Fuel-03

Steam-03

Fuel-07

Steam/demi-water tank Demi-water-05

DEMI WATER

SANITARY WATER

Sanitary water-01

Cracking

-1 97 kg h-1 345 kg h

8328 kg h-1 70°C

V-101

Steam-02

Carck-01

P-103

EXHAUST GASES

Ex-09

Demi-water-04

Fuel-05

SANITARY WATER

248 kg h-1

Ex-04

Steam-04

8328 kg h-1 25°C

Demi-water-02

S-101

P-101

E-109

Ex-03

Fuel-06

Fuel-04

Fuel-01 DIESEL

E-103

E-102 Ex-02

Demi-water-03

Fuel-02

Ex-08

Steam-01

E-101

E-104

T-101

APS

Air-04

b

FB-101

Air-06

Air-05

S-102

Ex-01

Cath-04

An-01

Cath-06

Anode

FCS 294 kg h-1

CB-101

Cath-05

Ref-03

P-102

F-101 Air-07

DEMI WATER

E-107

E-108

Ref-04

Demi-water-01

77 kg h-1

An-02

Cathode

Ref-05

Ref-02

371 kg h-1

Cath-02

Cath-01 An-03

An-04

Demi-water-00

E-106

V-102

Cath-03

Demi-water-05

Ref-01

Air-01

AIR

Steam/demi-water tank

ATR

Steam-04

K-101

Fuel-03

P-103

T-101 Cath-07

2040

international journal of hydrogen energy 34 (2009) 2026–2042

hull service, safety service, cargo service, engine room service, air conditioning and ventilation, gallery service, accommodation service and light service. The power requirement was estimated in sea-going conditions, manoeuvring operations, cargo operations and emergency. As far as the main auxiliary plants are concerned, the following were considered for comparison with our APU requirements: demi-water producers and diesel oil feeding facilities. Table 6 collects the above mentioned information, referred to specific reference cases of cruise ferries, Ro–Ro ships and Ro-Pax ships [56]. The total power requirement was lumped in one single value, expressed as the maximum attainable one in pessimistic conditions. The most relevant contributions came from: i) Cruise ferry: air conditioning and ventilation, engine room and cargo services, whose maximum required power was attained in sea-going conditions; ii) Ro–Ro ship: deck and hull service, for manoeuvring operations; iii) Ro-Pax ship: air conditioning and ventilation service and cargo service, in cargo handling operations. It can be observed that a suitable number of APU modules of 500 kWe, driven in steady-state at the full-load nominal operating conditions, could supply an important fraction of the power requirement coming from auxiliaries. The advantage is both in terms of efficiency compared to power generation from diesel engines, and environmental footprint. If one focuses on the raw materials, it arises that the on board rough consumption of fuel for auxiliaries is far greater than the amount needed for the 500 kWe APU (100 kg h1 for the best configuration in Fig. 6b); this implies that the facilities to stock and distribute diesel are already sized for the implementation of this system on board. Instead, the hourly capability of the demi-water generators (included water for sanitary purposes)

Table 6 – Utility capabilities already present on-board of reference types of ship. Max Demi-water Diesel auxiliaries production feeding capacity power capacity [kg h1] requirement [kg h1] [kW] Cruise Ferry Gross tonnage: 36,000 ton Cargo: 2800 crew þ passengers, 700 cars Ro–Ro ship Gross tonnage: 55,000 ton Cargo: 2500 lanemetres Ro-Pax ship Gross tonnage: 45,000 ton Cargo: 4200 lanemetres, 500 passengers

3330

1458

4300

1250

928

5000

1250

720

2900

is not large enough to withstand the additional water makeup required by the APU system without water recovery (345 kg h1 from Fig. 6b). It comes out that specific facilities should be installed on board to sustain the water requirements of the new system. On the grounds of this statement, the modification of the APU system in order to recover water (demi-water consumption reduced to 97 kg h1) could minimize the ship plant adjustments. In addition, within the demiwater on board ships requirements in Table 6, the fraction devoted to sanitary hot water production could be profitably integrated within the MCFC system in CHP version, since the amount of sensible heat available for this operation was able to provide at best 8328 kg h1 of hot sanitary water (scheme of Fig. 8b).

5.

Conclusions

A system modelling was carried out to compare different schemes for the MCFC technology for APU purposes in the marine field. The reference plant had a power of 500 kWe provided by the FC stack, plus roughly 100 kWe from a gas turbine. Two systems were compared in order to assess their efficiency: these systems differed in the way the fuel (marine diesel JP5) was reformed and processed to the FC stack. The first one was based on an ATR: the main advantages were that the ATR was almost thermal-neutral and that the huge amount of CO could be directly fed to the MCFC stack with no further cleaning steps, except for gas desulphurization. The efficiency of such configuration was 32.7%. The flue gases could be exploited to pre-heat the ATR feed, even if this reduced the flexibility of the system: with such degree of thermal integration the efficiency increased up to 39.1%. The second system was based on two steps: firstly, the liquid fuel was cracked into a gaseous fraction and a liquid residue, then the gaseous stream was conveyed to a steam reformer. Steam reforming is a very energy consuming step, however in the MCFC vessel a high amount of heat is produced (which must be removed to maintain a constant temperature into the stack) in the FC for the H2/CO reactions, and in the catalytic burner where the unreacted H2/CO is completely oxidized. Therefore, this configuration is extremely effective in exploiting an amount of heat that would have been wasted, as appears in Table 4 for the ‘‘ATR’’ based system. The efficiency of the ‘‘Cracking þ SR’’ based system was 44.6%. If the flue gases, following the same approach mentioned for the ATR-based system, were further integrated into the process, the efficiency attained 50.6%. The described schemes were upgraded to CHP purposes: the excess of sensible heat of the flue gases could be used to produce sanitary water. Thus, a CHP efficiency could be defined as the sum of the electrical efficiency (which remained unchanged with respect to the above mentioned values), and the one due to heat recovery. The CHP efficiency for the ‘‘ATR’’ based system reached 74.1% while the one referred to the ‘‘Cracking þ SR’’ system achieved the considerable and very attractive value of 85.6%. Finally, demineralised water could be recycled within the process, by condensing the steam of the flue gases. In this way

international journal of hydrogen energy 34 (2009) 2026–2042

both fixed and operating costs were reduced, and the CHP system was environmentally more efficient. The present results were compared to the power needs and raw materials on-board stock capabilities of three reference cases of commercial ships (cruise ferries, Ro–Ro and Ro-Pax vessels). Considering the size of the selected ships, these could profitably benefit of high efficiency MCFCs for APU purposes, without major ship plant adjustments for feeding the required diesel oil, while demi-water needs to be recycled within the APU to make the latter less impacting on the ship infrastructure.

Acknowledgments The support of CETENA S.p.A. (http://www.cetena.it/), and particularly of Ing. Marco Schembri, is gratefully acknowledged.

references

[1] U.S. Congress, Office of Technology Assessment. Marine applications for fuel cell technology – a technical memorandum. OTA-TM-O-37. Washington, DC: U.S. Government Printing Office; February 1986. [2] Hua J, Wu YH, Jin PF. Prospects for renewable energy for seaborne transportation – Taiwan example. Renew Energy 2008;33:1056–63. [3] Weaver G, Barrett S. Marine applications of fuel cell technology. Fuel Cells Bull 2003;1:11–2. [4] Monanteras NC, Nakos LG, Frangopoulos CA. Marine applications of fuel cells – an overview. Int Assoc Mech Eng Trans 2004;1(2):330–6. [5] Adamson K. Fuel cell today market survey: niche transport (part 2), APU and marine applications. London: Fuel Cell Today; 2006. [6] European Commission. Quantification of emissions from ships associated with ship movements between ports in the European Community. Entec UK Limited. Available from: http://ec.europa.eu/environment/air/; 2002. [7] European Commission. A European Union strategy to reduce atmospheric emissions from sea-going ships. COM:595. Available from: www.europa.eu.int/comm/environment/air; 2002. [8] European Commission. Fuel cells in ships – synthesis of open problems and roadmap for future RTD. Available from: www. europa.eu.int/comm/environment/air; 2004. [9] Natural Resources Defense Council. Harboring pollution: strategies to clean up U.S. ports. New York, NY. Available from: http://www.nrdc.org/air/pollution/ports; 2004. [10] Ahmed S, Krumpelt N. Hydrogen from hydrocarbon fuels for fuel cells. Int J Hydrogen Energy 2001;26:291–301. [11] Dincer I. Technical, environmental and exergetic aspects of hydrogen energy systems. Int J Hydrogen Energy 2002;27:265–85. [12] Dunn S. Hydrogen futures: toward a sustainable energy system. Int J Hydrogen Energy 2002;27:235–64. [13] Veldhuis IJS, Richardson RN, Stone HBJ. Hydrogen fuel in a marine environment. Int J Hydrogen Energy 2007;32:2553–66. [14] Ewald R. Requirements for advanced mobile storage systems. Int J Hydrogen Energy 1998;23(1):803–14. [15] Mori D, Hirose K. Recent challenges of hydrogen storage technologies for fuel cell vehicles. Int J Hydrogen Energy, in press doi:10.1016/j.ijhydene.2008.07.115. [16] Giacomazzi G. Prospects for intercontinental seaborne transportation of hydrogen. Int J Hydrogen Energy 1989;14(8): 603–16.

2041

[17] Adams VW. Possible fuel cell applications for ships and submarines. J Power Sources 1990;29(1–2):181–92. [18] Schmal D, Kluiters CE, Barendregt IP. Testing of a De Nora polymer electrolyte fuel cell stack of 1 kW for naval applications. J Power Sources 1996;61(1–2):255–7. [19] Sattler G. PEFCs for naval ships and submarines: many tasks, one solution. J Power Sources 1998;71(1–2):144–9. [20] Sattler G. Fuel cells going on-board. J Power Sources 2000;86: 61–7. [21] Beckhaus P, Dokupil M, Heinzel A, Souzani S, Spitta C. Onboard fuel cell power supply for sailing yachts. J Power Sources 2005;45:639–43. [22] Kickulies M. Fuel cell power for maritime applications. Fuel Cells Bull 2005;9:12–5. [23] Figueroa RA, Otahal J. Utility experience with a 250 kW MCFC cogeneration power plant at NAS Miramar, San Diego. J Power Sources 1998;71:100–4. [24] Aicher T, Lenz B, Gschnell F, Groos U, Federici F, Caprile L, et al. Fuel processors for fuel cell APU applications. J Power Sources 2006;154:503–8. [25] Krummrich S, Tuinstra B, Kraaij GJ, Roes J, Olgun H. Diesel fuel processing for fuel cells – DESIRE. J Power Sources 2006; 160:500–4. [26] Sarioglan A, Olgun H, Baranak M, Ersoz A, Atakul H, Ozdogan S. Diesel evaporation as the first step of hydrogen production. Int J Hydrogen Energy 2007;32:2895–901. [27] Tsourapas V, Sun J, Nickens A. Modeling and dynamics of an auto-thermal JP5 fuel reformer for marine fuel cell applications. Energy 2008;33(2):300–10. [28] Marine fuel cell market analysis. Washington, DC: USCG Research & Development Center; September 1999. CG-D-01-00. [29] Raugei M, Bargigli S, Ulgiati S. A multi-criteria life cycle assessment of molten carbonate fuel cells (MCFC) – a comparison to natural gas turbine. Int J Hydrogen Energy 2005;30:123–30. [30] Alkaner S, Zhou P. A comparative study on life cycle analysis of molten carbonate fuel cells and diesel engines for marine application. J Power Sources 2006;158:188–99. [31] U.S. Coast Guard Research and Development Center, U.S. Department of Transportation. Marine molten carbonate fuel cell demonstration module: USCGC VINDICATOR ship interface studies. CG-D-12-99; 1999. [32] METHAPU prototypes methanol SOFC for ships. Fuel Cells Bull 2008;5:4–5. [33] Navy contract helps PCI develop reformer tech. Fuel Cells Bull 2008;4:8. [34] Proton for world’s first fuel cell passenger ship. Fuel Cells Bull 2008;2:3. [35] HotModule power for FellowSHIP project. Fuel Cells Bull 2008;1:4. [36] Europe using fuel cells to clean up shipping. Fuel Cells Bull 2007;10:4. [37] H2 logic to power icelandic ship. Fuel Cells Bull 2007;8:10. [38] Proton power working on fuel cell ship. Fuel Cells Bull 2007; 3:8. [39] Tennessee research focuses on fuel cell applications for ‘e-ships’. Fuel Cells Bull 2006;12:3. [40] FCE extends naval contract for shipboard fuel cells. Fuel Cells Bull 2006;9:1. [41] MTI Micro wins Marines, Army DMFC contracts. Fuel Cells Bull 2004;12:4. [42] Power þ Energy studying use of diesel in marine fuel cells. Fuel Cells Bull 2004;10:6–7. [43] Hydrogenics module for marine application. Fuel Cells Bull 2004;6:6. [44] PEM fuel cell power for yachts. Fuel Cells Bull 2004;3:7. [45] Marra D, Bosio B. Process analysis of 1 MW MCFC plant. Int J Hydrogen Energy 2007;32:809–18.

2042

international journal of hydrogen energy 34 (2009) 2026–2042

[46] Specchia S, Saracco G, Specchia V. Modeling of an APU system based on MCFC. Int J Hydrogen Energy 2008;33:3393–401. [47] Detail specification grades JP-4 and JP-5. Available from: www.wbdg.org/ccb/FEDMIL/dtl5624u.pdf; 2004. [48] Kim JH, Ma X, Zhou A, Song C. Ultra-deep desulfurization and denitrogenation of diesel fuel by selective adsorption over three different adsorbents: a study on adsorptive selectivity and mechanism. Catal Today 2006;111:74–83. [49] Bhandari VM, Ko CH, Park JG, Han SS, Cho SH, Kim JN. Desulphurization of diesel using ion-exchanged zeolites. Chem Eng Sci 2006;61:2599–608. [50] Edenhofer R, Lucka K, Kohne H. Low temperature oxidation of diesel–air mixtures at atmospheric pressure. Proc Comb Inst 2007;31:2947–54.

[51] Perry RH, Green DW. Perry’s chemical engineers’ handbook. 7th ed. New York: McGraw Hill; 1988. [52] Cutillo A, Specchia S, Antonini M, Saracco G, Specchia V. Diesel fuel processor for PEM fuel cells: two possible alternatives (ATR versus SR). J Power Sources 2006;154: 379–85. [53] Rosso I, Galletti C, Bizzi M, Saracco G, Specchia V. Zinc oxide sorbents for the removal of hydrogen sulfide from syngas. Ind Eng Chem Res 2003;41:1668–97. [54] Ansaldo Fuel Cells S.p.A. private communication; 2007. [55] Campbell TJ, Shaaban AH, Holcomb FH, Salavani R, Binder M. JP-8 catalytic cracking for compact fuel processors. J Power Sources 2004;129:81–9. [56] CETENA S.p.A. private communication; 2008.