Ali Hedayati

Olivier Le Corre

Bruno Lacarrière

Dynamic Jordi Llorca

• Conference contributions

Keywords: Membrane reactor, ethanol steam reforming, pure hydrogen, hydrogen yield, hydrogen recovery, exergy analysis, exergy efficiency, thermal efficiency, static model, dynamic simulation, Sieverts' law

Production in-situ d'hydrogène pur par reformage d'éthanol dans un réacteur catalytique à membrane

My supervisors, Professor Jordi Llorca at Universitat Politècnica de Catalunya, Professor Olivier Le Corre and Professor Bruno Lacarrière at Ecole des Mines de Nantes, for giving me the opportunity to work with the interesting subject of pure hydrogen production in a catalytic membrane reactor. I greatly appreciate your patience, availability, openness and the quality of your advice and comments to my various questions. I remain ever grateful to you.

I would like to thank Professor Ignasi Casanova and Professor Bruno Lacarrière for all your constructive advice. Special thanks for organizing the opportunity for me to work at Universitat Politècnica de Catalunya and Ecole des Mines de Nantes via SELECT+ program.

Dr. Albert Casanovas, Dr. Lluís Soler, Dr. Núria J. Divins, and Dr. Adrian Miguel Schifer, for your friendly help. Without your guidance, especially during difficult times, I would have not been able to put this work together.

Thanks to all the people in the Institute of Energy Technologies (INTE) in Universitat Politècnica de Catalunya for contributing with your great spirit: Núria, Albert, Lluís, Raúl, Alejandra, Ander, Adrian, David, Raquel, Lourdes, Sara, Dani, and Pablo. My appreciation also goes to all the people at the Institute of Energy Technologies (INTE) • Proposer un modèle intégré dynamique du réacteur pour un usage ultérieur de contrôle dans la chaîne complète réacteur-pile à combustible. L'hydrogène produit (après membrane) a pour objectif d'alimenter une pile à combustible fournissant des besoins électriques domestiques par exemple. Ce système de production fonctionne en instationnaire. Or, la production d'hydrogène ne varie pas linéairement avec les paramètres de commande (précédemment désignés par paramètres opératoires). Des modèles statiques, correspondant aux gains de fonctions de transfert, basés sur des lois d'Arrhénius ont été développés pour les différentes espèces (H 2 , CO, CO 2 , CH [START_REF]NASA polynomials[END_REF] ). Pour la membrane, le modèle de Sievert a été ajusté et comparé avec une approche « boite noire ». Ensuite un modèle dynamique isotherme a été établi sur la base du premier principe et en supposant que l'équation du gaz parfait était acceptable.

Remarque : Dans ce manuscrit, le terme « carburant » est un abus de langage communément présent dans la littérature et désigne le mélange éthanol et eau.

Quelques éléments de contexte

Les engagements mondiaux sont résolument orientés vers une diminution des émissions anthropiques de CO 2 : diminution de la consommation et substitution de la consommation d'énergie fossile par des ressources moins ou peu émettrices de CO 2 .

Pour sa première promotion 2012-2016, le programme EMJD SELECT+ s'intéresse aux applications résidentielles, représentant 11% des émissions de CO 2 dans l'Union Européenne, voir Fig. 2. [START_REF] Capros | EU Energy, Transport and GHG Emissions Trending to 2050-Rerefence Scenario[END_REF] Le concept de « maison à basse (voire très basse) consommation d'énergie » ou celui de « maison à énergie positive » sont en progression. La production locale d'énergie est donc en développement, plus particulièrement la production d'énergie électrique à partir de sources renouvelables. La problématique de ces ressources se trouve au niveau de l'intermittence et en conséquence du stockage. L'une des manières de

Fig. 2. Emission de CO 2 par secteur dans l'Union Européenne

Dans le chapitre 2, un état de l'art de la production d'hydrogène par voie catalytique est proposé.

Production H 2 : résultats expérimentaux

Dans le chapitre 3, un banc d'essai de • le facteur de production d'hydrogène théorique (dans le réacteur, i.e. amont membrane), Hydrogen Yield noté :

= . × (5)
le débit H 2 récupéré (après membrane . ) par rapport au débit H 2 produit . . • La part de H 2 récupérée sur l'hydrogène produit

. par les réactions catalytiques, noté :

= . . (6)
• L'efficacité de production de l'hydrogène est définie comme l'hydrogène après membrane (le produit utile) sur l'hydrogène introduit dans le carburant . = . . [START_REF] Bernstein | Climate Change 2007: An Assessment of the Intergovernmental Panel on Climate Change[END_REF] Sur la Fig. 3, le facteur de production théorique H 2 après membrane est tracée. Plus la pression dans le réacteur est importante, plus le facteur de production H 2 croit, dans la gamme étudiée. De même, plus la température dans le réacteur croit, plus le facteur de production H 2 croit : une linéarité semble correctement décrire la tendance observée. La composition du carburant, caractérisée par le facteur SC est plus difficile à analyser car un effet croisé avec la température apparaît. En outre, le facteur de production H 2 ne croit pas avec le débit de carburant. molaire H 2 , CO, CO 2 , CH 4 et H 2 O. Le tableau 3 propose les valeurs des constantes du modèle à 5 coefficients. specie α specie [mol.m -3 ] specie [-] θ specie [J.mol -1 .Pa -1 ] specie [J.mol -

= δ - ( � , -� , ) (23)
Avec l'épaisseur de la membrane, , la pression après la membrane (prise à la presion atmosphérique, , la pression partielle de l'hydrogène dans le réacteur, une énergie d'activation, supposée constante et le facteur pré-exponentiel.

, = × , (24)
Le facteur pré-exponentiel peut être obtenu à partir des mesures :

= ×δ - ×(� , -� , ) (25)
Et est ensuite modélisé par :

= × × ( × ) (26)
Dont les coefficients et sont ensuite obtenus par la méthode des moindres carrés, voir tableau 5.

La Fig. 17 donne le facteur pré-exponentiel mesuré vs molélisé.

Il est maintenant possible de modéliser le modèle dynamique iso-therme : Le réacteur est décomposé comme :

-Un volume : application de la loi du gaz parfait et bilan de masse -Une zone catalytique : utilisation du modèle d'Arrhénius ajusté -Une membrane : utilisation du modèle de Sieverts ajusté Une constante de temps de la zone catalytique est ajoutée.

Le bilan de masse, entrée -sortie, du réacteur est :

= ṁ -ṁ -ṁ . (27)
L'équation d'état du gaz présent dans le volume du réacteur :

= ( ) × (28)
La dynamique de la zone catalytique-membrane est représentée par un modèle du premier ordre:

= + (29)
est le débit obtenu par la modélisation statique (régime permanent) et la constante de temps de la zone catalytique-membrane.

La vanne actionnant le débit de carburant étant considérée comme « immédiate », le modèle de cet actionneur est :

= ( 30
)
La dynamique de la zone catalytique-membrane est :

= + (31)
La Fig. 19 donne la description du modèle simulink réalisé.

List of figures

Fig. 1.1 CO 2 emission by sector in EU [START_REF] Capros | EU Energy, Transport and GHG Emissions Trending to 2050-Rerefence Scenario[END_REF] Fig. 1.2 Electricity generation by fuel type in EU [START_REF] Capros | EU Energy, Transport and GHG Emissions Trending to 2050-Rerefence Scenario[END_REF] Fig. 2.1 Scheme of alcohol reforming using a catalytic membrane reactor for generating hydrogen (B) compared to a conventional system based on in-series reactors (A).

List of tables

Personal reference

• Journal publications

Chapter 1

Introduction and general overview

Abstract

In this chapter, a general overview of the environmental concerns of the world energy demand and power production is given. The importance of clean energy production and the solutions for on-site energy supply especially in the case of the buildings are discussed briefly. Finally, the objectives and scope of this thesis together with the outlines are presented.

Background and general overview

It is proven that the increasing rate of energy consumption and high dependency of energy production on fossil fuels has caused obvious negative environmental consequences of climate change. According to the intergovernmental panel on climate control (IPCC), growing concentration of greenhouse gases in the atmosphere -mostly by burning of fossil fuels for power generation processes -is the main reason of climate change and global warming [START_REF]Intergovernmental Panel on Climate Change IPCC[END_REF]. Svante Arrhenius was one of the first who mentioned global warming probability due to the increasing concentration of CO 2 in the atmosphere [START_REF] Arrhenius | On the Influence of Carbonic Acid in the Air upon the Temperature of the Ground[END_REF]. He suggested that the mean temperature of the earth would probably increase due to the emission of carbon dioxide originating from human activities. Now the atmospheric concentrations of CO 2 have increased from natural mean value of 280 ppm to 379 ppm in 2005 and it is estimated to exceed 400 ppm by 2030 [START_REF] Bernstein | Climate Change 2007: An Assessment of the Intergovernmental Panel on Climate Change[END_REF]. It is expected that the worldwide energy consumption will be increasing rapidly so that the yearly CO 2 emissions from energy production sectors will reach 33.8 Gt in 2020 and 42.4 Gt in 2035 [8]. According to the IPCC, a 50-85% reduction in total CO 2 emission by 2050 is mandated to limit the anticipated rise in global temperature to within β˚C [START_REF]IPCC Fourth Assessment Report: Climate Change 2007 (AR4)[END_REF].

Besides, energy conversion is a key factor for the development of human society so that living without energy supply -mainly electricity -is not possible in the modern world. Thus, as development continues with an aim to eradicate poverty and enhance quality of life, energy consumption by all the sectors of the society is unavoidable [10,[START_REF] Klugman | Human Development Report[END_REF]. A rapid shift towards efficient and cost effective sources of energy is obviously required. These facts have forced human to start serious movements towards clean energy production and energy saving strategies.

Among different sectors of the society, it is indicated that in the modern world more than 40% of the total produced energy is demanded by buildings, mainly in the form of electricity [START_REF] Kapsalaki | A methodology for economic efficient design of Net Zero Energy Buildings[END_REF]. Consequently, a big share of emission of greenhouse gases can be attributed to the buildings so that according to the reference scenario 2013 by European Union, 11% of total CO 2 emission belongs to only residential buildings [START_REF] Capros | EU Energy, Transport and GHG Emissions Trending to 2050-Rerefence Scenario[END_REF]. European electricity demand is estimated to increase from 3081 TWh in 2012 to 3250 TWh in 2020 [13]. The CO 2 emission by sector is given in Fig. 1.1. CO 2 emission from buildings originates from on-site energy generation and burning fuels for heat in buildings or cooking in homes. CO 2 emission from electricity use in buildings is included in the Power generation/District heating sector. Renewable energy resources are now considered as the fastest and most feasible solution to achieve the targets of clean electricity production. Nevertheless, one of the most important challenges is the dependency of availability and utilization of renewable sources on the geographical and local conditions and infrastructures. Besides, inevitable CO 2 emission and most importantly, transmission of produced electricity to the end users remain among the challenges to be encountered. In this regard, on-site electricity production to lower the dependency of different sectors of the society on gridded power sources has been investigated recently. Hence, new solutions such as small scale on-site energy generation units in buildings may open a new era to harness available clean energy sources efficiently at the place/time where needed.

The concept of Zero Energy Building (ZEB) has recently received significant attention as a realistic solution to electricity and power generation for the building sector applications independent from gridded power. Although Zero Energy Buildings (ZEB) concept does not have a commonly agreed and clear definition, it is in general referred to a building which is able to provide itself with the required energy demand; in any form needed (i.e. electricity, heating, cooling, lighting. etc) [START_REF] Marszal | Zero Energy Building -A review of definitions and calculation methodologies[END_REF][START_REF] Sartori | Net zero energy buildings: A consistent definition framework[END_REF]. A sub-definition of ZEB is the theory of Net Zero Energy Building (NZEB) denoting a ZEB, which is connected to the grid. In this case, the building is capable of supplying the grid (when the building produces energy more than it needs) with the same amount of energy that it has obtained from the grid (when the energy produced by the building is lesser than the requirements). According to Marszal [START_REF] Marszal | Zero Energy Building -A review of definitions and calculation methodologies[END_REF] and Sartori [START_REF] Sartori | Net zero energy buildings: A consistent definition framework[END_REF], the quantity of energy taken out from the grid, is equal to the quantity which is sent back to the grid by the building, considering normally a duration of one year. Moreover, if a building total energy production over a specific period exceeds its needs and requirements, so that the excess energy is given to the grid, it is defined as a Positive Energy Building.

Currently, serious worldwide attempts are under way by the researchers and policy makers for development and commercialization of ZEBs. EU policies indicate that buildings should be nearly Net Zero Energy Building by 2020 [START_REF] Kapsalaki | A methodology for economic efficient design of Net Zero Energy Buildings[END_REF] and also the US Department of Energy (DOE) has the strategy of achieving commercial ZEBs by 2025 [START_REF]Building Technologies Program;Planned Program Activities for[END_REF]. In view of the explanations of ZEBs, the idea of on-site production of clean energy is a key factor for ZEBs to develop. Although renewable sources seem the first option, as mentioned before, issues such as economical and feasibility aspects of solar and wind energy for small scale (home scale) applications, and the dependency of these resources upon geographical and local climate conditions must be taken into account.

Among different methods and technologies tested for small scale on-site electricity production -mainly to be used in the buildings and light transportation sector -, fuel cells technology has shown proven potentials for different tasks and can be applied in sub-MW size at any condition independent from local factors. Being compatible with modern energy carriers such as hydrogen, fuel cells are considered as the energy convertors of the future power generation systems [START_REF] Sharaf | An overview of fuel cell technology: Fundamentals and applications[END_REF]. In a fuel cell, chemical energy of a fuel (mainly hydrogen) is converted into electricity via a chemical reaction of positively charged hydrogen ions with a source of oxygen, which is air. The products of such a conversion are electricity, water, and heat. To summarize, the advantages of fuel cells can be listed as:

• Reduced harmful emissions: in the case of utilizing pure hydrogen as fuel, the emission (of the fuel cell stack) is supposed as zero (excluding the emissions for the H 2 production). This is considered as the most important advantage of fuel cell technology over the heat engines and conventional power cycles. • High efficiency: up to 60% efficiency has been reached, and in the case of high temperature fuel cells, considerable amount of heat can be recovered. • Modularity: the power output of the fuel cell can be adjusted easily by adding/removal of the stacks. • Prompt load flowing: thanks to the instant nature of the electrochemical reactions, the dynamic behavior of fuel cell in responding to the load change is reliable. • No noise pollution: fuel cells are static apparatuses and due to absence of mechanical moving parts, no noise pollution is caused. • Fuel flexibility: not only the pure hydrogen, but also some carbonaceous fuels such as light alcohols are directly used in the fuel cells as the source of hydrogen.

Pure hydrogen is used as the main fuel for fuel cells to produce electrical power (plus lots of recoverable heat from high temperature operating types) via electrochemical reactions. While the combustion of hydrogen produces very clean energy, its conversion to electricity and heat by means of fuel cells is known as the highest efficient energy producing process [START_REF] Koch | Ethanol catalytic membrane reformer for direct PEM FC feeding[END_REF]. The high efficiency of the fuel cells is attributed to the single-step energy conversion pathway during which chemical energy is converted to electrical energy [START_REF] Sharaf | An overview of fuel cell technology: Fundamentals and applications[END_REF]. This is comparable to the complicated conventional power generation cycles where chemical energy is converted to thermal energy, then to the mechanical energy and finally to the electrical energy considering the thermodynamic losses occurring in each process.

Apart from pure hydrogen, several research works has been focused on direct application of methanol, ethanol, and ethylene glycol mainly for prevention of pure hydrogen production units. A comprehensive review on the fuel cells has been published by Sharaf and Orhan [START_REF] Sharaf | An overview of fuel cell technology: Fundamentals and applications[END_REF]. The case of Direct Ethanol Fuel Cells (DEFC) is briefly discussed due to the similarities between this concept and the concept of this PhD work. The necessity of pure hydrogen production can be a matter of question while some liquid fuels can be directly used in fuel cells. The idea of DEFC was first developed and investigated in 1995 when ethanol steam reforming in a Molten Carbonate Fuel Cell (MOFC) was simulated and studied [START_REF] Freni | Ethanol steam reforming in a molten carbonate fuel cell: a thermodynamic approach[END_REF]. DEFCs are one type of the Direct Alcohol Fuel Cells (DAFSc) in which a type of alcohol -mainly methanol or ethanol -is used directly in the electrochemical process [START_REF] Kamarudin | Review: Direct ethanol fuel cells[END_REF]. DEFCs operate at very high temperatures with efficiency between 20% and 40%, however, the development and technology approval of this kind of fuel cell stays at early stages [START_REF] Sharaf | An overview of fuel cell technology: Fundamentals and applications[END_REF].

Some positive points such as ease of transport and delivery of ethanol, and higher energy density of ethanol compared to hydrogen, cause DEFSc to compete in efficiency and price with Proton Exchange Membrane Fuel Cells (PEMFCs, as the most commercial type of fuel cells for car and building applications). However, some challenges are hindering the commercialization of DEFSc. Very high sensitivity to carbon monoxide, low performance in C-C bond breakage due to low temperature (considered as the main feedback), acetic acid formation and inability of the oxidation of this specie, and permeation of ethanol through the membrane are the most important challenges to be conquer for the optimization and further development of DEFCs [START_REF] Abdullah | Modeling and simulation of a direct ethanol fuel cell: An overview[END_REF]. According to Leone et al. [START_REF] Leone | Operation of a solid oxide fuel cell under direct internal reforming of liquid fuels[END_REF], DEFCs operating at 700˚C and higher still suffer from carbon deposition. Additional water in the feed may hinder carbon formation to some extent, but on the other hand lowers the current density of the fuel cell. As the open literature presents, there is not a lot of study performed on the development and investigation of DEFCs, although since 2007 the amount of research work has increased notably. A comprehensive literature review on DEFCs was presented in the work by Kamarudin et al. [START_REF] Kamarudin | Review: Direct ethanol fuel cells[END_REF].

According to the discussion above, the application of DEFCs is not technically approved and has a long way to be optimized and developed to enter the market and compete the commercial and well-developed types of fuel cells -mainly PEMFCs.

Although fuel cell technology is among the cleanest techniques to produce energy, the main challenge remains unresolved. How to produce and deliver pure hydrogen in a sustainable, safe, applicable, and efficient way? Conversion technologies for pure hydrogen production and most importantly the saving and transportation of pure hydrogen are still a concern. To be economically feasible and comparable, hydrogen needs to be liquefied or compressed (high risk of high pressure), or adsorbed in metallic hydrides in large scale prior to the transfer (very low temperature or huge masses of metallic hydrides). Hence, it can be concluded that it is highly beneficial to produce and consume pure hydrogen at the same place/time.

Hydrogen can be obtained by different techniques such as electrolysis of water, catalytic reforming of alcohols, catalytic conversion of methane, photochemical splitting of water, and biological conversion of waste biomass. To reach the cleanest way, it is inevitable to convert a renewable source to hydrogen.

Ethanol itself as a beverage has been in used for thousands of years and has been produced via fermentation of natural resources. However, this especial organic chemical has been in use broadly for various applications such as a solvent, an antifreezing agent, disinfecting agent, and most importantly as a fuel. Ethanol is a hydrogen rich fuel with hydrogen to carbon ratio of 3. Besides, there are ethanol distribution infrastructures in many countries. Ethanol is not considered as a toxic liquid, is safe to reserve at ambient temperature and pressure, not explosive, and its special molecular structure makes it easier for C-C breakage in reforming processes.

Bioethanol is referred to the ethanol produced via biological conversion of biomass (fermentation or enzymatic catalysts) [START_REF] Llorca | Hydrogen from bio-ethanol[END_REF] and is also applied for the alcoholic liquid products obtained by the mentioned fermentation process. The molar alcohol to water ratio of such products is between 1:7 and 1:12. If ethanol is considered as a fuel, further distillation for water removal is necessary which means highly energy consuming and costly processes. Instead, the mixture of water and ethanol in the form of bioethanol directly can be used as a feedstock for reforming processes. This leads to significant reduction in global energy consumption and the costs of the process by omitting the additional alcohol purification unit operations [START_REF] Deluga | Renewable hydrogen from ethanol by autothermal reforming[END_REF]. Depending on the availability of the biomass, the source of hydrogen differs from region to region through the world. For example, sugar cane is used in Brazil and grains like corn are used in US. In 2012, Brazil and US as the largest ethanol producers, produced around 86% of ethanol used globally [24]. In Europe, the leaders of bio-ethanol production are Germany, Spain, France and Sweden, mostly from beet and wheat [START_REF] Markets | Creating Markets for Renewable Energy Technologies EU RES Technology Marketing Campaign[END_REF]. The price of the edible plants as a source of fuel ethanol has increased recently resulting in a conflict between using them as a source of fuel or a source of food. Besides, the increasing trend of the universal need of food has resulted in recent researches in the use of non-edible natural resources such as grasses and straw to produce bioethanol, although the bioethanol production process from these kind of resources is more complicated due to the complex molecular structures [START_REF] Llorca | Hydrogen from bio-ethanol[END_REF].

Noble metal-based catalysts are well known for very high reactivity in terms of ethanol conversion and hydrogen selectivity together with nearly zero carbon deposition over the surface of the catalyst [START_REF] Llorca | Hydrogen from bio-ethanol[END_REF][START_REF] Tosti | Design and process study of Pd membrane reactors[END_REF][START_REF] Gallucci | Pd-Ag membrane reactor for steam reforming reactions: A comparison between different fuels[END_REF]. Among noble metals, Rh, Pt, Pd, Ru, and Ir have been in the center of attention during last 10 years [START_REF] Contreras | Catalysts for H2 production using the ethanol steam reforming (a review)[END_REF]. Ethanol steam reforming (ESR) over noble metal-based catalysts can be considered as an efficient and reliable method for hydrogen production. The application of membrane reactors (MR) -in which production and separation of hydrogen (pure hydrogen production) occur in the same reactor vessel -is highly beneficial to omit costly and complicated unit processes for hydrogen purification. Besides, by removal of one of the products (hydrogen) via permeation through the membrane, equilibrium limitations are overcome even at unbeneficial operating conditions, leading to higher production of hydrogen and higher efficiency of the reforming system [START_REF] López | Hydrogen production from ethanol over Pd-Rh/CeO2 with a metallic membrane reactor[END_REF].

In the case of a palladium-based membrane, fuel cell grade hydrogen can be obtained, so that a fuel cell can be fed online as pure hydrogen is generated (permeated).

The objectives of the PhD work

The main objective of this thesis is to perform a comprehensive investigation on a pure hydrogen generating system -named as the reformer -in which ethanol steam reforming is performed over noble metal-based catalysts in a membrane reactor for on-site production of fuel cell grade hydrogen. The specific scopes are:

• To perform ethanol steam reforming (ESR) over Pd-Rh/CeO 2 in a membrane reactor (MR) containing Pd-Ag membranes to observe the performance of the MR in the ESR atmosphere at several different operating conditions aiming to introduce the best operating conditions in terms of hydrogen production and recovery o Experimental study of the performance of a Pd-Rh catalyst in ethanol steam reforming (ESR) atmosphere at different operating conditions regarding temperature, pressure, flow rate of the ethanol-water mixture (fuel), and different molar ratios between ethanol and water (S/C ratio), in terms of ethanol conversion, hydrogen and byproduct selectivity, the permeation rate of hydrogen through the membrane, hydrogen yield and hydrogen recovery.

o Comparison between the CMR (presented in this thesis) and the Staged Membrane Reactor configuration (SMR) [START_REF] López | Hydrogen production from ethanol over Pd-Rh/CeO2 with a metallic membrane reactor[END_REF][START_REF] Koch | Ethanol catalytic membrane reformer for direct PEM FC feeding[END_REF] reported in the literature to emphasize the enhancements and benefits of the CMR.

• To perform a thermodynamic analysis on the performance of the reformer at different operating conditions in terms of energy and exergy efficiency to understand the most favorable operating conditions based on the exergy efficiency, and to obtain detailed information on the sources of thermodynamic losses and to introduce the possible ways to optimize the reforming system thermodynamically.

o Exergy evaluation of the reformer based on the experimental results. o Investigation of the sources of exergy loss and exergy destruction. o Analysis of the heat losses and potential resources of recoverable exergy to give relevant solutions to increase the exergy efficiency of the reformer aiming to optimization of its thermal performance.

o Comparison between the CMR (presented in this thesis) and the Staged Membrane Reactor configuration (SMR) [START_REF] López | Hydrogen production from ethanol over Pd-Rh/CeO2 with a metallic membrane reactor[END_REF][START_REF] Koch | Ethanol catalytic membrane reformer for direct PEM FC feeding[END_REF] in terms of exergy evaluation factors based on the experimental results.

• To model and simulate the reformer presented in this thesis at steady and transient states regarding the production rate of ESR products and the permeation rate of pure hydrogen aiming to model a system capable of regulating the pure hydrogen production rate according to the modifications needed to adjust the electrical load of a fuel cell, when fed online.

o Simulation of the production rate of different species -especially hydrogen -as the main products of the ESR process in the catalytic membrane reactor (CMR) at tested operating conditions.

o Simulation of pure hydrogen production both based on the Sieverts' Law and a black box model, as an essential step for dynamic modeling of the reformer.

o Simulation of the dynamic performance of the reformer in terms of pure hydrogen production rate and its variations in the case of pressure or fuel flow rate set point modifications

Outline of the thesis

In the second chapter, an introduction to hydrogen energy is given followed by a comprehensive literature survey on the production of hydrogen via catalytic reforming of ethanol. Moreover, the state of the art is given regarding the catalytic steam reforming of ethanol in membrane reactors.

In An introduction to the hydrogen production via catalytic reforming processes, and the membrane reactors, together with a comprehensive literature review on the hydrogen production over common catalysts in the membrane reactors are given in the next chapter.

44 Chapter 2

Hydrogen production and membrane reactors

Abstract

The previous chapter provided a general overview and a short history of the world energy demand and probable solutions for clean energy production, together with the objectives and scope of this thesis. In this chapter, an introduction to the hydrogen production is given. A short history about using hydrogen as an energy carrier together with the current considerations of hydrogen as a renewable source of energy is presented. The introduction and definition of a membrane reactor (MR) is followed by the review of the production of hydrogen via catalytic steam reforming of ethanol in membrane reactors (MRs). The state of the art on the production of hydrogen in catalytic membrane reactors (CMRs) via steam reforming of ethanol over noble metal based catalysts is presented.

Hydrogen

The first ideas of using hydrogen as a source of energy appeared and revived around two centuries ago, when for the first time the French writer Jules Verne talked about the future when humankind will heat himself using water splitted into hydrogen and oxygen [START_REF] Goltsov | Biosphere synergism and the humankind virtual path to the hydrogen civilization era[END_REF][START_REF] Verne | Twenty thousand leagues under the sea[END_REF].

In the 20 th century, the idea of hydrogen energy was brought to the scientific and engineering level and developed rapidly by many research groups worldwide and finally in 1974 The Clean Energy Research Institute of Miami University (USA) established the International Association for Hydrogen Energy (IAHE). Soon after, the IAHE started to publish the International Journal of Hydrogen Energy (IJHE), which is one of the most well-known references of the scientific research literature on hydrogen-related research works.

Concerning the attempts to achieve the clean energy production targets, hydrogen as an energy carrier can serve the energy production sector with the necessary requirements. In short, the advantages of hydrogen as energy carrier over other energy sources are [START_REF] Cipriani | Perspective on hydrogen energy carrier and its automotive applications[END_REF]:

• If hydrogen is combusted with air, heat is released and pure water is produced as the only product; hydrogen is the only carbon-free fuel available for combustion engines. • It is produced via conversion of biofuels -for example bioalcohols -as promising replacements to the fossil fuels. • It is used in centralized or remote applications with reduced impact of energy production. • Thanks to commercially available fuel cells, hydrogen is the main fuel of fuel cells, as electricity producer devices with numerous advantages. • In the future, on-site hydrogen production can be possible (transportation infrastructure not needed).

However, there are challenges and obstacles to be overcome for the utilization of hydrogen as a reliable energy carrier for comprehensive applications in the society. Although a lot of researches have been done recently to investigate the techniques and methods of production of hydrogen, no authoritative position compared to other clean energy resources is seen yet in the energy policy frameworks [START_REF] Bleischwitz | Policies for the transition towards a hydrogen economy: the EU case[END_REF]. The European energy policy framework does not ignore hydrogen energy systems as carbonless sources of energy. But on the other hand, it has been sufficed to mention the importance of regulations and infrastructures development for hydrogen applications in the energy sector [START_REF] Commission | Comunication from the commission to the European parliament, the council, the European economic and social committee and the committee of the region[END_REF]. This can be attributed to some issues. First, hydrogen and fuel cells are in the early stage of entering the market commercially. Second, requirements of special infrastructures for distribution of hydrogen makes it complicated to utilize hydrogen and fuel cells. In conclusion, more supportive policies are needed to push hydrogen technology and systems toward an existing option for the energy sector.

Regarding the attempts for introducing the hydrogen energy systems, the concept of hydrogen economy has recently defined. Hydrogen economy in general aims to serve the human with hydrogen for its energy demand, considering the point that hydrogen is an energy carrier, not a source of energy [START_REF] Luis | Renewable Hydrogen Technologies[END_REF]. The same challenges facing the production of hydrogen will face to the hydrogen economy. There is no point in investments on hydrogen infrastructures for distribution and refueling, if there is no market for hydrogen. Again, the high cost of fuel cells and hydrogen production from renewable resources adds another challenge to the future of hydrogen economy. Nevertheless, according to McDowall and Eames [START_REF] Mcdowall | Forecasts , Scenarios , Visions , Bachcasts and Roadmaps to the Hydrogen Economy μ A Review of the Hydrogen Futures δiterature for UK-SHEC[END_REF] there are 4 major drivers of a hydrogen economy: climate change, energy security, local air quality, and competitiveness.

In the light of EU energy policies and by knowing the fact of taxation exemption of hydrogen energy, the effort to introduce hydrogen as a competitive and reliable energy carrier must be continued [START_REF] Bleischwitz | Policies for the transition towards a hydrogen economy: the EU case[END_REF]. The key factor to success in hydrogen technology development is to receive the governmental supports via energy policies set mainly by the two main powers i.e. EU and U.S.A. The economic, technological and institutional challenges of introduction of hydrogen in the energy systems in Europe are addressed by The European Hydrogen Energy Roadmap, known as HyWays [START_REF]HYWAYS the European Hydrogen Roadmap[END_REF]. The HyWays aims to open an eye and give ideas on how to receive the political support and overcome the barriers that hydrogen energy encounters. According to HyWays, 80% of the light duty vehicles and city buses should be supplied by clean hydrogen by 2050. Consequently, several initiatives such as 'Fuel Cells and Hydrogen Joint Technology Initiative' and 'Hydrogen and Fuel Cell Program Plan' were started in Europe and U.S.A, respectively [START_REF] Luis | Renewable Hydrogen Technologies[END_REF].

Hydrogen Production

As previously stated, the production of hydrogen is dependent on the technological advancements, policymaking, and even the price of fossil fuels as the main sources of energy. Apart from this, according to Ros et al. [START_REF] Ros | Snapshots of hydrogen uptake in the future A comparison study[END_REF] at least 75% of the produced hydrogen will be used in the transportation sector. Until 2030, the main hydrogen production technologies are anticipated to be methane steam reforming and electrolysis of water, both accompanied with carbon capture and sequestration (CCS) processes. But figures for 2050 perspective show that the hydrogen production would be mainly dependent on natural gas reforming and biomass conversion [START_REF]Prospects for Hydrogen and Fuel Cells[END_REF]. Again, the hydrogen production is claimed to be supported by CCS techniques. In this regard, and having in mind that CCS technology is not approved for the large scale and worldwide scale applications, the concept of CO 2 -free hydrogen production rises. Although renewable energy sources such as solar and wind can be utilized for production of electricity needed for electrolysis of water [START_REF] Ros | Snapshots of hydrogen uptake in the future A comparison study[END_REF], the challenge of renewable resources still exist. Electrolysis of water will require huge amount of electricity, which makes it unavoidable to be supplied by a clean electricity source, and on the other hand, natural gas reforming will result in enormous amount of CO 2 production that needs to be supported by CCS.

In general, the so-called renewable hydrogen can be produced by electrolysis or thermolysis of water (as long as the required electricity is supplied by wind or solar power plants), thermochemical and biological conversion of biomass, and photolysis of water [START_REF] Turner | Renewable hydrogen production[END_REF]. However, all these technologies are in the development and research stage and need to access renewably produced electricity.

Catalytic conversion of ethanol

Hydrogen is directly produced from ethanol by three main catalytic conversion routes i.e. steam reforming (ESR), partial oxidation (POX), and autothermal reforming (ATR). The main difference between the three ways refers to the converting coreactants used in each process. The pros and cons of each process are briefly discussed.

The highest amount of hydrogen is obtained by steam reforming of ethanol so that theoretically (complete conversion of coreactants to carbon dioxide and hydrogen) 6 moles of hydrogen is formed per one mole of ethanol in the feed. Ethanol can catalytically be converted to hydrogen and carbon dioxide in the presence of steam via the following reactions [24]: [START_REF] Domínguez | Ethanol steam reforming at very low temperature over cobalt talc in a membrane reactor[END_REF][START_REF] Divins | Bio-ethanol steam reforming and autothermal reforming in 3-m channels coated with RhPd/CeOβ for hydrogen generation[END_REF][START_REF] Men | Ethanol Steam Reforming in a Microchannel Reactor[END_REF]. The stoichiometric molar steam to ethanol ratio (S/E) is 3; however, higher amounts can be used. The bioethanol produced by fermentation of biomass has S/E ratio of 7-12 that can directly be used for the ESR process.

C 2 H 5 OH → H 2 + CO + CH 4 2.1 CO + H 2 O ⇆ H 2 + CO 2 2.2 CH 4 + 2H 2 O ⇆ 4H 2 + CO 2 2.3 -------------------------------------------- CH 3 CH 2 OH + 3H 2 O ↔ 2CO 2 +
The main drawback of ESR is the endothermicity. In one hand high energy input is needed for running the reaction and evaporating the liquid fuel, and on the other hand, the hydrogen formation is limited by thermodynamic equilibrium due to the nature of the reactions. The thermodynamic equilibrium limit can be overcome at S/E ratios higher than 20, leading to 5.5 mole hydrogen per mole of ethanol at 773-873 K [START_REF] Llorca | Hydrogen from bio-ethanol[END_REF].

At any rate, much higher energy input is needed at high S/E ratios for evaporation of water. Another disadvantage of this process is the formation of CO, which can poison the electrocatalysts of the PEMFCs and makes it inevitable to purify the produced hydrogen prior to feed the fuel cell. Water gas shift reaction promotion is one of the most investigated methods for CO removal together with preferential oxidation of CO.

In partial oxidation of ethanol (POX), ethanol is reacting with oxygen for production of hydrogen and carbon dioxide. Thus, the disadvantages of ESR are overcome but the hydrogen yield drops to three moles of hydrogen per mol ethanol (eq. 2.5).

CH 3 CH 2 OH + O 2 → 2CO 2 + 3H 2 2.
5

(ΔH 298 K = -552 kJ/mol)
The advantages of POX process over ESR can be summarized as [START_REF] Llorca | Hydrogen from bio-ethanol[END_REF]:

• POX is a self-sustaining process and no external heating is required. Consequently, not only the energy consumption is decreased, but also the process plan is simplified and more compact. • There is no equilibrium limitation resulting in operating at temperatures as low as 473 K. • The chemical reactions are faster and the reactor scale needed for POX process is smaller than the one needed for ESR.

Apart from this, oxygen is a gaseous coreactant that does not need evaporation, and the presence of oxygen hinders the carbonaceous species to be formed and the lifetime of the catalyst is extended.

The drawbacks of the POX lie on the nature of oxidizing process in presence of oxygen because the total oxidation (combustion) of ethanol and hydrogen is thermodynamically favored. Both reactions (combustion of ethanol and hydrogen) are highly exothermic and competing with the partial oxidation of ethanol. This may cause low hydrogen selectivity. Besides, the mixture of ethanol and oxygen at high temperature in gas phase raises the risk of explosion.

ATR reaction is a combination of ESR and POX so that the reforming process is performed in presence of water steam and oxygen aiming to utilize the heat released by POX for ESR. For this reason, the concentration of oxygen and steam must be adjusted to be near the stoichiometric values to take the advantage of a thermally neutral process. Also by correctly adjustment of oxygen content in the fuel, the disadvantages of POX can also be avoided and the process can be operated at lower temperatures rather than ESR. The presence of water steam can decrease the risk of explosion and the competitiveness of combustion reactions.

The disadvantage of ATR is the complexity of the process and the difficulty of adjustments for a steady state operation.

In general, the energy demand per mole of hydrogen produced during reforming processes follows the trend of ESR ≥ ATR > POX. The highest hydrogen yield belongs to the ESR, despite being endothermic. As a result, major share of research has been devoted to the investigation and technological maturity of ESR.

Side products of reforming processes

The main reaction pathway of ethanol reforming starts with the decomposition of ethanol to two important intermediates i.e. acetaldehyde and ethylene according to the following reactions:

CH 3 CH 2 OH → CH 3 CHO + H 2 2.6 CH 3 CH 2 OH → C 2 H 4 + H 2 2.7
The presence of these two species has been reported by some researchers [START_REF] Vaidya | Kinetics of Steam Reforming of Ethanol over a Ru / Al 2 O 3 Catalyst[END_REF][START_REF] Lu | Reaction network and kinetic analysis of ethanol steam reforming over a Ru/Al2O3 catalyst[END_REF]. Ethylene and acetaldehyde are formed prior to formation of carbon oxides. Ethylene plays an important role in coke formation, which is one the most important factors leading to catalyst deactivation. On the other hand, acetaldehyde further promotes the formation of hydrogen. Both intermediates undergo steam reforming resulting in the formation of carbon oxides, hydrogen, methane and ethane [START_REF] Llorca | Hydrogen from bio-ethanol[END_REF][START_REF] Lu | Reaction network and kinetic analysis of ethanol steam reforming over a Ru/Al2O3 catalyst[END_REF]. Lu et al. [START_REF] Lu | Reaction network and kinetic analysis of ethanol steam reforming over a Ru/Al2O3 catalyst[END_REF] reported the observation of ethane formation that could be a result of ethylene hydrogenation:

C 2 H 4 + H 2 → C 2 H 6 2.8 (ΔH 298 K = 14 kJ/mole)
The reactions in presence of water are endothermic (except WGS) and lead to more hydrogen production. Methane generation mechanisms are exothermic and are favored at low temperatures leading to less hydrogen production. The formation of coke in reforming processes is common [START_REF] Palma | On the activity of bimetallic catalysts for ethanol steam reforming[END_REF]. Coke formation is a result of several complex reactions. In this regard, the formation of ethylene has an important role of the deposition of carbon on the surface of the catalyst:

C 2 H 4 → polymeric chains → 2C + 2H 2 2.9
According to the open literature, numerous catalysts have been investigated for the ESR process and, among them noble metals have shown superior catalytic reactivity in terms of higher hydrogen production (higher ethanol conversion) and robustness of the performance (durability) at different operating conditions [START_REF] Ni | A review on reforming bio-ethanol for hydrogen production[END_REF][START_REF] Lopez | Ethanol Steam Reforming Thermally Coupled with Fuel Combustion in a Parallel Plate Reactor[END_REF]. Further, the formation of undesired chemical species is almost zero when noble metal-based catalysts are used for ESR process. This is considered as one of the main advantages of noble metals over other catalysts that may compensate to some extent the relatively higher price of these catalysts compared to conventional ones.

Ethanol steam reforming in the literature

A catalyst must maintain special properties to be appropriate for the ESR process. The most important characteristics of catalysts can be named as ability to dissociate the C-C band preferably at low temperature, high reactivity toward CO conversion, and durability under catalytic reaction conditions [START_REF] Llorca | Hydrogen from bio-ethanol[END_REF]. Before 2007, Rh and Ni were the most commonly tested catalysts for hydrogen production via ESR [START_REF] Ni | A review on reforming bio-ethanol for hydrogen production[END_REF][START_REF] Haryanto | Current status of hydrogen production techniques by steam reforming of ethanol: A review[END_REF]. After 2007, Co and other noble metals such as Pd, Pt, and Ir, and bimetallic active phases attracted a lot of attention due to enhanced reactivity and robustness in the ESR environment. The catalytic activity depends both on the selection of the relevant support and the preparation method. Thanks to special characteristics that favor ethanol dehydrogenation but inhibit dehydration, supports such as MgO, ZnO, CeO 2 , La 2 O 3 , promoted Al 2 O 3 , and promoted ZrO 2 were introduced as suitable supports for Ni and Ru in terms of hydrogen production and long-term stability [START_REF] Contreras | Catalysts for H2 production using the ethanol steam reforming (a review)[END_REF].

Additionally, the development of bimetallic catalysts results in higher hydrogen production and catalyst stability. The promoted ethanol conversion and hydrogen selectivity in the case of the development of bimetallic catalysts have been reported by several researchers [START_REF] Koh | Highly efficient ruthenium and ruthenium-platinum cluster-derived nanocatalysts for hydrogen production via ethanol steam reforming[END_REF][START_REF] Andonova | Structure and redox properties of Co promoted Ni/Al2O3 catalysts for oxidative steam reforming of ethanol[END_REF][START_REF] Shi | La2O2CO3 supported Ni-Fe catalysts for hydrogen production from steam reforming of ethanol[END_REF][START_REF] Scott | Hydrogen production from ethanol over Rh-Pd/CeO2 catalysts[END_REF][START_REF] Su Lee | Hydrogen-rich gas production from ethanol steam-reforming reaction using NiZr-loaded MCM-48 catalysts at mild temperature[END_REF]. Investigations on the performance and applicability of the catalysts for ethanol steam reforming can be divided into two groups: catalysts without noble metals, and catalysts containing noble metals.

Ethanol Steam Reforming over catalysts without noble metals

Cobalt and nickel are among the most tested catalysts for ESR process. Cobalt is a cheap catalyst able to convert ethanol at temperatures as low as 623 K so that WGS reaction is enhanced and CO concentration is kept low. Besides, methane is not formed as an intermediate, leading to higher hydrogen yield. Co/Mg/Al hydrocalcite catalysts over ceramic honeycombs were tested by Espinal et al. [START_REF] Espinal | Cobalt hydrotalcite for the steam reforming of ethanol with scarce carbon production[END_REF] for ESR and it was reported that the hydrocalcite with a Co:Mg:Al molar ratio of 1:2:1 showed the best catalytic performance with very small coke formation. Dominguez et al. [START_REF] Domínguez | Co-Fe-Si Aerogel Catalytic Honeycombs for Low Temperature Ethanol Steam Reforming[END_REF] used cobalt talc doped with iron for ESR process and reported that at 673 K and 2 bar, 1.04 l N H 2 per ml of ethanol is obtained at steam to carbon ratio of 3 (S/C=3). In another study, Espinal et al. [START_REF] Espinal | Cobalt hydrotalcites as catalysts for bioethanol steam reforming. The promoting effect of potassium on catalyst activity and long-term stability[END_REF] studied the effect of the addition of potassium to catalytic regarding the activity and stability of Co/Mg/Al hydrocalcite catalysts for ESR process. Very low carbon deposition together with presence of oxidized cobalt as an active species was observed. As mentioned by Llorca [START_REF] Espinal | Ethanol steam reforming in a catalitic membrane reactor loaded with a hydrotalcitederived cobalt catalyst[END_REF], cobalt catalyst doped over K + was tested successfully for bio-ethanol conversion during 300 hours without carbon formation, which is an advancement in ethanol conversion over cobalt-based catalyst. Dominguez et al. [START_REF] Domínguez | Ethanol steam reforming at very low temperature over cobalt talc in a membrane reactor[END_REF] tested a catalytic membrane reactor consisting of cobalt talc (Co 3 [Si 2 O 5 ] 2 (OH) 2 ) supported on cordierite. The experiments were performed in a Pd-Ag metallic membrane to separate hydrogen from produced gases. The test was done in two ways. First, honeycomb catalysts were placed in series, and second, the membrane was covered by small pieces of the catalyst and a superior reactivity in terms of pure hydrogen production was reported.

Nickel-based catalysts are used due to their known and approved high capability of C-C bond cleavage [START_REF] Murakami | Metal-catalysed cleavage of carbon-carbon bonds[END_REF]. Ni is normally supported on alumina for better endurance in the reaction conditions but still having the disadvantage of being prone to carbon deposition [START_REF] Llorca | Hydrogen from bio-ethanol[END_REF]. Consequently, other elements are added to Ni-based catalysts as a promoter to the catalytic activity. Supports such as MgO, SiO The high activity was ascribed to the high dispersion of Ni and intimate contact between the metals, which resulted in carbon formation prevention on the catalyst surface [START_REF] Yang | Production of hydrogen by steam reforming of ethanol over a Ni/ZnO catalyst[END_REF]. Souza et al [START_REF] De Souza | Synthesis Gas Production by Steam Reforming of Ethanol over M-Ni-Al Hydrotalcite-type Catalysts; M=Mg, Zn[END_REF] studies Co, Mo, and Zn prepared in Ni-based hydrotalcite catalyst at 673-873 K and concluded that lower reaction temperatures favored the dehydration and dehydrogenation of ethanol, while synthesis gas was produced mainly at higher temperatures . Co-, Mo-and unmodified samples presented higher selectivity towards H 2 and CO at higher temperatures, while samples modified with Co, Mo and Zn showed higher activity at intermediate temperatures [START_REF] De Souza | Synthesis Gas Production by Steam Reforming of Ethanol over M-Ni-Al Hydrotalcite-type Catalysts; M=Mg, Zn[END_REF].

Ethanol Steam Reforming over noble metal-based catalysts

As discussed before, one of the major problems of catalytic steam reforming of ethanol is carbon deposition on the catalyst leading to the severe deactivation. It is proven that noble metals not only are able to highly convert ethanol due to their significant reactivity, but also hinder carbon from deposition on the active sites [START_REF] Lopez | Ethanol Steam Reforming Thermally Coupled with Fuel Combustion in a Parallel Plate Reactor[END_REF][START_REF] Idriss | A phenomenological study of the metal-oxide interface: the role of catalysis in hydrogen production from renewable resources[END_REF]. This distinctive property has attracted the attention of a lot of research groups toward noble metals for ESR application so that according to Llorca et al. [START_REF] Llorca | Hydrogen from bio-ethanol[END_REF] nearly 40% of the published works talked about Rhbased catalysts and other 30% corresponds to the evaluation of Pt-based ones.

According to the literature, CeO [START_REF] Palma | CeO2-supported Pt/Ni catalyst for the renewable and clean H2 production via ethanol steam reforming[END_REF]. The study was done at 523-873 K and very high selectivity of catalyst was reported leading to very low coke formation even at stoichiometric ratio of ethanol/water. The catalyst selectivity improved by increasing the amount of water in the feeding fuel.

Goula et al. [START_REF] Goula | Hydrogen production by ethanol steam reforming over a commercial Pd/ -Al2O3 catalyst[END_REF] used commercial Pd/ -Al 2 O 3 for production of hydrogen rich gas via ESR and they obtained hydrogen selectivity up to 95% at 873 K. Complete ethanol conversion reached at even low temperature and the selectivity hydrogen and CO 2 decreased with S/C ratio. The same catalyst was tested by Auprêtre et al. [START_REF] Auprêtre | Bio-ethanol catalytic steam reforming over supported metal catalysts[END_REF] and hydrogen selectivity of 55% was reached at 973 K and atmospheric pressure under stoichiometric reaction conditions (S/C≈1.5). Chen et al. [START_REF] Meng | Hydrogen Production from Steam Reforming of Ethanol over Pd Promoted ZnO/Al2 O3 Catalysts[END_REF] studied Pd/ZnO-Al 2 O 3 at 723 K and concluded that acetaldehyde and hydrogen were the main product of ESR reactions and coke formation was not hindered by Pd promotion. Ethanol conversion of 65% together with hydrogen selectivity of 65% were reached, which are considerably lower compared to similar reported studies. Other authors mentioned that Pd/CeO 2 -Y 2 O 3 stabilized by ZrO 2 showed good initial activity but significant deactivation was observed [START_REF] Ramos | Hydrogen production from ethanol steam reforming on M/CeO2/YSZ (M=Ru, Pd, Ag) nanocomposites[END_REF]. Complete ethanol conversion was reached at 873 K where hydrogen selectivity was 67-74%. Ru/Pd/Ag supported on CeO 2 /YSZ was used for ESR experiments and the Ru-based one showed the superior activity. Positive effect of CeO 2 on preventing the carbon deposition plus metal selectivity of catalyst to syngas was mentioned as the main reason of the good performance of the mentioned catalyst. Scott et al. [START_REF] Scott | Hydrogen production from ethanol over Rh-Pd/CeO2 catalysts[END_REF] found that the bimetallic catalyst 0.5%Rh-0.5%Pd is able to break the C-C bond of ethanol at 400 K. This catalyst showed maximum ethanol conversion (nearly complete) and hydrogen selectivity (64%) at 773 K. Pd/HZSM-5 catalyst with 5%wt Pd and prepared by wet impregnation method was tested by Lie et al. [START_REF] Wang | Steam Reforming of Model Compounds from Bio-Oil for Hydrogen Production over Pd/HZSM-5 Catalyst[END_REF]. The best performance was reached at 973 K and S/C=9.2, where the hydrogen selectivity was 58.1-84.3%. Frusteri et al. [START_REF] Frusteri | H2 production for MC fuel cell by steam reforming of ethanol over MgO supported Pd, Rh, Ni and Co catalysts[END_REF] reported the selectivity of 70% for Pd/MgO. The selectivity towards CO 2 decreased drastically because of catalyst deactivation, which consequently increased the selectivity towards CO. Acetaldehyde, and ethylene were both detected (15% and 6%, respectively). The study of the temperature effect on the performance of Pd-Ru/Nb 2 O 5 -TiO 2 in terms of hydrogen production in the ESR environment showed the maximum performance at 573 K [START_REF] Alonso | Temperature Effect on Hydrogen Production from Reactions between Ethanol and Steam in the Presence of Pd-Ru/Nb2O5-TiO2 Catalyst[END_REF]. The presence of H catalyst was reported by Auprêtre et al. [START_REF] Auprêtre | Bio-ethanol catalytic steam reforming over supported metal catalysts[END_REF]. The activity of Ni-based catalyst was compared to Ru-based one and it was reported that although Ni-based catalyst gave higher hydrogen yield, its selectivity towards CO 2 was lower. In the case of Ce x Zr 1-x O 2 mixed oxide-supported 1 wt.%Rh catalysts, the best performance in terms of the highest hydrogen yield and the lowest coke deposition belonged to the Rh/Ce 0.5 Zr 0.5 O 2 at 873 K [START_REF] Birot | Ethanol steam reforming over Rh/CexZr1-xOβ catalystsμ Impact of the CO-CO2-CH4 interconversion reactions on the H2 production[END_REF]. At all operating conditions (temperature and S/C ratio); the ethanol steam reforming yielded a large amount of methane, which reduced the hydrogen production. The formation of methane was ascribed to the CO hydrogenation. Roh et al. [START_REF] Roh | Low temperature and H2 selective catalysts for ethanol steam reforming[END_REF] focused on a support prepared by a co-precipitation method having composition of Ce Idriss et al. [START_REF] Idriss | A phenomenological study of the metal-oxide interface: the role of catalysis in hydrogen production from renewable resources[END_REF] showed that Rhodium in presence of Platinum or palladium is able to completely convert ethanol to gaseous species. Besides, they reported very stable activity with no sign of carbon deposition after three weeks of operation for ESR process. The reported results clearly represent the advantages of noble metal-based catalysts and the possibility of overcoming the major challenges of ESR process i.e. coke formation, and catalytic durability. The robustness of noble metals under the operational conditions can be considered as compensation to the high price of these materials.

Regarding the application of ceria, still, better results belong to the combination of Rh and Pd supported on ceria. It has been proved that high oxygen storage capacity (OSC) together with the properties such as easy reducibility via facile Ce III ↔ Ce IV equilibrium, makes ceria a very durable material at high temperatures and very active during reducing/oxidizing processes [START_REF] Sundararajan | Cooperative Synergy in Nanoscale Ceria-Based Systems[END_REF].

Catalytic Membrane Reactors (CMRs)

The reaction temperature is an important operational parameter when conducting alcohol and bio-alcohol reforming reactions. High temperatures are necessary for bond cleavage, but moderate temperatures are preferred for the WGS equilibrium to favor the formation of hydrogen and CO 2 at the expense of CO and water, thus maximizing the production of H 2 and avoiding the requirement of bulky WGS units at the reactor outlet. For that reason, the use of catalytic membrane reactors (CMRs), where the generation and separation of hydrogen take place simultaneously (extractor membrane), appears as a very attractive approach to strongly simplify alcohol reformers. In addition, the shift effect that occurs in CMRs results in even higher hydrogen yields because the presence of a membrane selective to the hydrogen permits attaining very high conversion values in comparison with the traditional reactors operating under the same conditions [START_REF] Gallucci | Pd-Ag membrane reactor for steam reforming reactions: A comparison between different fuels[END_REF]. In fact, the continuous removal of one of the reaction products, the hydrogen, promotes the reaction conversion beyond the equilibrium values. With respect to a classical configuration consisting of a reactor unit in series with a separation unit, CMRs represent a modern configuration in which an integrated reaction/separation unit has many potential advantages: reduced capital costs, improved yields and selectivities and drastically reduced downstream separation costs [START_REF] Basile | Hydrogen Production Using Pd-based Membrane Reactors for Fuel Cells[END_REF][START_REF] Mendes | Integrated analysis of a membrane-based process for hydrogen production from ethanol steam reforming[END_REF]. In addition, the reject gas from the membrane can be used as a fuel source for a catalytic combustor to provide a self-sustainable operation [START_REF] Montané | Thermodynamic analysis of fuel processors based on catalytic-wall reactors and membrane systems for ethanol steam reforming[END_REF][START_REF] Hedayati | Exergetic study of catalytic steam reforming of bio-ethanol over Pd-Rh/CeO2 with hydrogen purification in a membrane reactor[END_REF].

If hydrogen is intended to feed a low-temperature fuel cell based on proton exchange membranes (PEMFC), it needs to be purified in addition to WGS by means of further processes like pressure swing adsoprtion (PSA), CO preferential oxidation (COPrOx) or membrane separation. In this context, membrane reactor technology plays an important role as an advanced solution to conventional systems based on in series reactors, offering the opportunity of combining the reforming process for producing hydrogen and its separation in the same device (Fig. 2.1). P H2 is the hydrogen permeability, p H2,r and p H2,perm are the hydrogen partial pressures in the retentate and permeate sides of the membrane, respectively, and δ is the membrane thickness. The thiner the membrane the higher the hydrogen flux. However, to ensure mechanical resistance and strenght of the membrane, thicknesses greater than 5 m are commonly required. Assuming that the solid-state diffusion of hydrogen is ratedetermining, and that hydrogen atoms form an ideal solution in the metal, the hydrogen pressure exponent (n) is equal to 0.5, which corresponds to the Sieverts' law:

J H2 =P H2 (p 0.5 H2,r -p 0.5 H2,perm )/ 2.11

Then, a relatively large difference in the partial pressure of hydrogen between the reaction and separation sides is necessary for promoting hydrogen separation. Deviations from the Sieverts' law (n>0.5) may be attributed to the accumulation of impurities on the membrane surface or pinholes or microcracks in the membranes, or it may be attributed to masstransport resistance. On the other hand, the hydrogen permeability dependence on the temperature can be expressed by a classical Arrhenius equation. Therefore, when the Sieverts' law is valid, the hydrogen flux can be expressed as:

J H2 =P 0 H2 [(exp(-E a /RT)](p 0.5 H2,r -p 0.5 H2,perm )/ 2.12 which is known as the Richardson's equation. P 0 H2 is the pre-exponential factor and E a the apparent activation energy. The hydrogen permeation through the membrane is very sensitive to E a .

As stated by Al-Mufachi et al. [START_REF] Al-Mufachi | Hydrogen selective membranes: A review of palladium-based dense metal membranes[END_REF] , a variety of commercial dense metal membranes such as vanadium (V), niobium (Nb) and tantalum (Ta) have been studied to have the highest hydrogen permeability. The major drawback of such metals is their tendency to form stable oxide layers under ambient conditions. Palladium-based membranes are favorable in this regard (Pd) showing relatively high hydrogen permeability and are the suitable alternatives as a dense metal membrane.

Pd-based membrane is able to dissociate molecular hydrogen into monatomic form ready for fast diffusion through its lattice. Besides, in 1866, Thomas Graham discovered that Pd is capable of absorbing approximately 600 times its own volume in hydrogen whilst maintaining its physical properties and structural integrity [START_REF] Al-Mufachi | Hydrogen selective membranes: A review of palladium-based dense metal membranes[END_REF]. Commercial applications of Pd-based permeators for producing very pure hydrogen have been studied for more than fifty years (i.e. in the fuel cycle of nuclear reactors). In fact, palladium membranes are among the oldest membranes studied for gas permeation and separation applications and are still the membranes with the highest hydrogen permeability and selectivity [START_REF] Gao | Chemical Stability and Its Improvement of Palladium-Based Metallic Membranes[END_REF]. They are receiving renewed attention because of the prospect of the hydrogen economy. PEMFC require high-purity hydrogen for operation, with less than ca. 20 ppm of CO. The most critical issue for practical applications of Pd membranes in CMRs for hydrogen production is the chemical stability of the metal membranes (poisoning effects of the reaction mixture on hydrogen permeation, carbon deposition on the membrane, etc.). The poisoning through exposure to hydrogen sulphide affects negatively Pd membranes, which can be destroyed irreversibly. This may be particularly important for reforming reaction of bio-alcohols, where H 2 S traps may be necessary. The presence of CO in the reformate can cause the decrease of the hydrogen permeation at low temperature because CO molecules adsorb on Pd and displace adsorbed hydrogen, blocking the hydrogen adsorption sites (competitive adsorption). Water also blocks hydrogen permeability through the adsorption of oxygen atoms that form on the Pd surface during recombinative desorption of adsorbed water molecules. Coke affects negatively the hydrogen permeation by covering the membrane surface and lowering the hydrogen permeating flux. In addition, carbon atoms penetrate into the Pd lattice causing membrane failure owing to the concomitant expansion of the Pd lattice.

For a better chemical stability (poisoning) and physical stability (mechanical stress, hydrogen embrittlement), membrane reactors do not use pure Pd but various different types of Pd alloyed with other metals such as silver, copper, nickel, iron and platinum. Most Pd alloy membranes studied are of binary components, with a few of multicomponents. Pd-Ag (23 wt% Ag) and Pd-Cu (38-42 wt% Cu) alloy membranes are nowadays widely employed in CMRs for hydrogen production. The Pd-Ag is a plastic alloy with a specific hydrogen permeability of P H2 =3.4 nm 3 •mm•m -2 •h -1 MPa -0.5 at 873 K [START_REF] Burkhanov | Palladium-Based Alloy Membranes for Separation of High Purity Hydrogen from Hydrogen-Containing Gas Mixtures[END_REF]. The permeation rate of hydrogen through the membrane is a function of the membrane properties, and also the driving force, which is given by the difference of the square root of the hydrogen partial pressure on each sides of the membrane (Sieverts' law). The hydrogen permeation via a dense Pd-based membrane follows the mechanism of solution-diffusion transport. The membrane thickness plays an important role in the permeation rate, but on the other hand the mechanical strenght of the membrane must be taken into account. Membranes are generally deposited onto porous supports such as SiO 2 , Al 2 O 3 , B 2 O 3 and porous stainless steel (PSS), the later being advantageous because it has mechanical durability, a thermal expansion coefficient close to that of Pd, and ease of gas sealing. Unfortunately, PSS alloys the Pd at high temperatures, leading to lowering of the hydrogen permeability. This is usually solved by using an intermediate inert layer.

Pd membranes have been prepared by a variety of methods such as cold-rolling, evaporation, sputtering, spray pyrolysis, chemical vapor deposition, electroplating, and electroless plating [START_REF] Uemiya | Brief Review of Steam Reforming Using a Metal Membrane Reactor[END_REF]. During the rolling, a dense film without any pinholes and cracks can be maintained to the level of the appropriate thickness because the film is mechanically formed from the corresponding dense metal lump [START_REF] Tosti | Diffusion bonding of pd-ag rolled membranes[END_REF]. In the other methods a thin film is formed by rebuilding metal atoms and clusters, so that a considerable effort is required to avoid pinholes. Via cold-rolling and diffusion welding, robust Pd-based thin wall tubes less than 0.05 mm wall thickness have been systematically produced and their complete hydrogen selectivity and durability have been demonstrated in long term tests of CMRs [START_REF] Tosti | Long-term tests of Pd-Ag thin wall permeator tube[END_REF]. Today, numerous catalytic membrane reactors designs are available containing a bundle of Pd-based tubes or fingers [START_REF] Tosti | Design and process study of Pd membrane reactors[END_REF][START_REF] Burkhanov | Palladium-Based Alloy Membranes for Separation of High Purity Hydrogen from Hydrogen-Containing Gas Mixtures[END_REF]. These devices are capable for producing high hydrogen throughputs and can be used in compact reforming systems.

Pd alloy membranes have been used in catalytic membrane reactors mainly for WGS and steam reforming reactions of methane and methanol [START_REF] Basile | Hydrogen Production Using Pd-based Membrane Reactors for Fuel Cells[END_REF], but their use in the steam reforming of ethanol and higher alcohols is relatively new [START_REF] Iulianelli | Hydrogen production from ethanol via inorganic membrane reactors technology: a review[END_REF]. The main aim of the researchers involved in this field is oriented to emphasize the role of the membrane by analyzing the performances of the reaction system in terms of alcohol conversion, hydrogen yield and hydrogen recovery, that is, the amount of hydrogen collected in the permeate side vs. the total hydrogen produced during the reaction.

Ethanol steam reforming in CMRs

A large piece of work on ethanol steam reforming (ESR) with Pd-Ag membranes (wall thickness of 50 m) using sweep gas has been carried out by Basile's group at the University of Calabria, Italy. Gallucci et al. [START_REF] Gallucci | Methanol and ethanol steam reforming in membrane reactors: An experimental study[END_REF] studied ethanol steam reforming over Ru/Al 2 O 3 catalyst in three different Pd-Ag membrane reactors. The highest conversion attained at 723 K was about 50% in counter-current mode, which was significantly higher than that attained in a traditional reactor. However, in the CMR coke deposition occurred to a large extent with consequent deactivation of the catalyst. Basile et al. [START_REF] Basile | CO-Free Hydrogen Production by Ethanol Steam Reforming in a Pd-Ag Membrane Reactor[END_REF] used the same catalyst for performing ESR in a dense Pd-Ag CMR by varying the water:ethanol molar ratio between 3:1 and 9:1 at 573-673 K and 1.3 bar in counter current mode. Hydrogen recovery values of 22% were reached as well as ethanol conversion higher than 99% with a 56% hydrogen yield, which represented a significant improvement over the performance obtained with conventional catalytic reactors according to the shift effect [START_REF] Basile | Hydrogen Production by Ethanol Steam Reforming: Experimental Study of a Pd-Ag Membrane Reactor and Traditional Reactor Behaviour[END_REF]. Over the same catalyst, Tosti et al. [START_REF] Tosti | Low temperature ethanol steam reforming in a Pd-Ag membrane reactorPart 1: Ru-based catalyst[END_REF] at ENEA (Italy) reached hydrogen yields as high as 80% working at 723 K and 2 bar using a dense thin wall Pd-Ag tube. The same device was used by the same authors [START_REF] Tosti | Lowtemperature ethanol steam reforming in a Pd-Ag membrane reactorPart 2. Pt-based and Ni-based catalysts and general comparison[END_REF] for studying Pt-and Ni-based catalysts, which showed a poor performance in terms of hydrogen yield with respect to the Ru-based catalyst (Ru>Ni>Pt at low feed, while for higher feed flow ratios the sequence was Ru>Ni=Pt).

Tosti et al. [START_REF] Tosti | Ethanol steam reforming kinetics of a Pd-Ag membrane reactor[END_REF] also obtained kinetic expressions and modeled the CMR operation with the Ru-, Ni-and Pt-based catalysts in order to optimize the membrane reformer by assessing the ratio between the reaction and permeation kinetics. The membrane tube was divided into finite volume elements where the mass balances for both lumen and shell sides are carried out. The model was validated via experiments using three different kinds of catalyst, i.e. Ru, Pt and Ni-based. The effect of pressure was studied by Tosti et al. [START_REF] Tosti | Pressure effect in ethanol steam reforming via dense Pd-based membranes[END_REF] in a Pd-Ag tube of 150 mm wall thickness in the range 1-8 bar. A Ru-based catalyst was used and water/ethanol mixtures of molar ratio 10/1 and flow rates of 5, 10 and 15 g h -1 was fed. At 723 K and 4 bar and under a feed flow rate of 5 g h -1 , maximum values of hydrogen yield (5.5) and hydrogen recovery values close to 100% were measured. At lower temperatures and pressures the hydrogen yield declined. Lulianelli et al. [START_REF] Iulianelli | Oxidative steam reforming of ethanol over Ru-Al2O3 catalyst in a dense Pd-Ag membrane reactor to produce hydrogen for PEM fuel cells[END_REF] studied the additon of oxygen to the ESR environment in a Pd-Ag membrane reactor. The Ru/Al 2 O 3 catalyst was studied at 673 K, S/C=5.5, GHSV=2000 h -1 and by using a sweep gas into the permeate side of the reactor, being the maximum hydrogen recovery (ca. 30%) achieved at O/C=1.2. It was claimed that oxygen addition can prevent ethylene and ethane formation caused by dehydration of ethanol as well as carbon deposition [START_REF] Iulianelli | Oxidative steam reforming of ethanol over Ru-Al2O3 catalyst in a dense Pd-Ag membrane reactor to produce hydrogen for PEM fuel cells[END_REF]. A combined methane and ethanol steam reforming was carried out over a Pt/Al 2 O 3 catalyst at 1-5 bar in a multimembrane module (Pd-Ag, 150 m), where pure H 2 was recovered in the shell side by vacuum pumping [START_REF] Borgognoni | Combined methane and ethanol reforming for pure hydrogen production through Pd-based membranes[END_REF]. The authors nicely showed that ethanol concentration could be avoided by the combined steam reforming of diluted ethanol (bio-ethanol) and methane, that is, the excess of water in the bio-ethanol is used to perform the methane reforming. At 5 bar, hydrogen yields up to 70% were obtained with a water:ethanol:methane molar ratio of 14:1:1.

ESR over Co/Al 2 O 3 catalyst was conducted by Lulianelli and Basile [START_REF] Iulianelli | An experimental study on bio-ethanol steam reforming in a catalytic membrane reactor. Part I: Temperature and sweep-gas flow configuration effects[END_REF] and Iulianelli et al. [START_REF] Iulianelli | An experimental study on bio-ethanol steam reforming in a catalytic membrane reactor. Part II: Reaction pressure, sweep factor and WHSV effects[END_REF]. Several operational parameters such as temperature, pressure, sweep gas flow (SF) and load were evaluated and hydrogen yield and recovery values as high as 60% and 95%, respectively, were reached at 673 K, 3 bar, and SF=25.2 mL•min -1 (countercurrent flow). Experiments were performed in two types of reactor, i.e. a traditional reactor (TR) and a membrane reactor (MR). 95% ethanol conversion was reached by the MR in both cocurrent and counter-current flow configurations, while only 84% was reached in the TR. The hydrogen selectivity was 67% and 64% in the case of MR and TR, respectively. [START_REF] Iulianelli | An experimental study on bio-ethanol steam reforming in a catalytic membrane reactor. Part I: Temperature and sweep-gas flow configuration effects[END_REF]. The effect of the reaction pressure and the sweep factor on the reaction system was studied by means of the counter-current sweep gas flow configuration [START_REF] Iulianelli | An experimental study on bio-ethanol steam reforming in a catalytic membrane reactor. Part II: Reaction pressure, sweep factor and WHSV effects[END_REF]. The same catalyst was tested at 67γ K in a porous stainless steel (PSS) supported Pd membrane reactor (β5 m Pd layer deposited onto a stainless steel tubular macroporous support) with the aim of investigating the influence of the membrane characteristics as well as of the reaction pressure from 3 to 8 bar [START_REF] Basile | Ethanol steam reforming reaction in a porous stainless steel supported palladium membrane reactor[END_REF]. 100% ethanol conversion was reached at 673 K and 8 bar. Hydrogen recovery of more than 50% with a purity around 65% was obtained. Co/Al 2 O 3 and Ni/ZrO 2 catalysts were used in a Pd/PSS membrane reactor at 673 K and 8-12 bar for simulating bio-ethanol steam reforming by using a mixture of water-ethanol-acetic acidglycerol with 1:13:0.18:0.04 molar ratio [START_REF] Seelam | Hydrogen production from bio-ethanol steam reforming reaction in a Pd/PSS membrane reactor[END_REF]. About 94% of ethanol conversion was obtained at 12 bar over the Co/Al 2 O 3 catalyst, with 40% hydrogen yield and 40% hydrogen recovery.

Papadias et al. [START_REF] Papadias | An analytical and experimental investigation of high-pressure catalytic steam reforming of ethanol in a hydrogen selective membrane reactor[END_REF] at Argonne National Laboratory, USA, explored the benefits of highpressure ESR for the production of hydrogen needed to refuel the high-pressure tanks of PEMFC vehicles. The experiments were conducted at 7-70 bar, 873-1023 K, and S/C=3-12 in a Pd-Ag (γ0 m) CεR loaded with Rh/δa 2 O 3 -Al 2 O 3 catalyst. As expected from thermodynamics, higher pressures showed inhibition of the hydrogen yield in favor of methane. Domínguez et al. from the Technical University of Catalonia, Spain, studied the ESR in a CMR over cobalt talc at 598-673 K and 5-15 bar [START_REF] Domínguez | Ethanol steam reforming at very low temperature over cobalt talc in a membrane reactor[END_REF]. In addition to an improvement of the hydrogen yield with respect to a fluidized bed reactor (FBR), the CMR showed a rapid response to changes in the ethanol-water mixture load; a constant hydrogen flow was obtained after 2 s following variations of the ethanol-water load of ±10%. The experiments of Papadias et al. [START_REF] Papadias | An analytical and experimental investigation of high-pressure catalytic steam reforming of ethanol in a hydrogen selective membrane reactor[END_REF] and Domínguez et al. [START_REF] Domínguez | Ethanol steam reforming at very low temperature over cobalt talc in a membrane reactor[END_REF] were performed without sweep gas, therefore, pure hydrogen was obtained in the permeate side of the membrane, ready to feed a PEMFC. López et al. [START_REF] López | Hydrogen production from ethanol over Pd-Rh/CeO2 with a metallic membrane reactor[END_REF] studied the ESR using a Pd-Rh/CeO 2 catalyst over cordierite monoliths in-series in a CMR. Reaction yields of 3.1 mol hydrogen generated per mol ethanol in feed and total yields of 1.4 mol H 2 permeated per mol ethanol in feed were measured, with maximum hydrogen recuperation of 70% (no sweep gas). This CMR was directly connected to a PEMFC and several control algorithms were developed to study suitable strategies for obtaining fast response from the CMR following variations in the hydrogen demand from the fuel cell [START_REF] Koch | Ethanol catalytic membrane reformer for direct PEM FC feeding[END_REF].

Iulianelli et al. [START_REF] Iulianelli | The oncoming energy vector: Hydrogen produced in Pd-composite membrane reactor via bioethanol reforming over Ni/CeO2 catalyst[END_REF] studied bioethanol steam reforming over A 7.5 wt.% Ni/CeO 2 catalyst in a thin layer Pd-based (8 m thin Pd-layer deposited onto a porous Al 2 O 3 support) membrane reactor at 673 K. The study focused on the effect of feed composition and reaction pressure (2-3 bar) on the performance of the membrane reactor in terms of hydrogen permeation rate, permeated hydrogen purity, and ethanol conversion. The best performance of the membrane reactor was reached the highest pressure (3 bar). Almost complete ethanol conversion and a hydrogen yield of 67% was reached. The evaluation of the effect of presence of bioethanol impurities such as glycerol (C 3 H 8 O 3 ) and acetic acid (CH 3 COOH) showed that these species could positively contribute to more hydrogen production via attending the steam reforming reactions, specially in presence of excess steam (high steam to carbon ratio) [START_REF] Iulianelli | The oncoming energy vector: Hydrogen produced in Pd-composite membrane reactor via bioethanol reforming over Ni/CeO2 catalyst[END_REF]. For the first time, self-heated carbon nanotubes (CNT) as membranes were fabricated and used in steam reforming of ethanol by Janas et al [START_REF] Janas | Steam reforming on reactive carbon nanotube membranes[END_REF]. The surface of CNT was used as both the support of a Ni catalyst to perform the ESR reaction, and the source of heat to run the ESR reactions. The main advantages of such a configuration are more uniform temperature distribution (heterogeneous product distribution due to the temperature gradients is prevented), and light and portable setups. The temperature of 723 K was reached where 25% of hydrogen was detected in the effluent stream. Murmura et al. [START_REF] Murmura | Directing selectivity of ethanol steam reforming in membrane reactors[END_REF] performed ethanol steam reforming over Pt/Ni-CeO 2 catalyst in a Pd-based membrane reactor (4-5 m palladium layer, deposited by electroless plating on the outside of a porous alumina tube). The experiments were performed at 613-753 K, 6-10 bar, steam to carbon ratio of 1.5 (stoichiometric conditions), and using a nitrogen sweep gas flow rate of 0.5 l N min -1 at the permeate side. At complete ethanol conversion, 4.5 mol hydrogen per mol EtOH in the feed was obtained at 753 K. The hydrogen recovery of 100% was reported in the case of using the sweep gas at the permeate side [START_REF] Murmura | Directing selectivity of ethanol steam reforming in membrane reactors[END_REF]. Rh/La 2 O 3 -SiO 2 and Rh/CeO 2 catalysts were studied in a self-supported Pd-Ag membrane reactor for ethanol steam reforming [START_REF] Da Silva | Study of the performance of Rh/La2O3-SiO2 and Rh/CeO2 catalysts for SR of ethanol in a conventional fixed-bed reactor and a membrane reactor[END_REF]. The reaction experiments were performed in a traditional reactor and a membrane reactor using different sweep gas flow rates to feed the reactor with the mixture of ethanol and water (the reactant mixture was 2.5% ethanol, 7.5% water, and 90% nitrogen). The only products of the ESR reactions were CO, CO 2 , CH 4 , and H 2 and complete ethanol conversion was kept in all cases. The reactions were favored in the case of the membrane reactor without coke formation. The best results were obtained at at S/C ratio of 10. Hydrogen recovery of 70% was reached where 2.8 mol hydrogen was produced per mol converted ethanol.

Although most of the ESR work in CMR has been done with Pd-Ag metallic membranes, there are also some examples of ESR in different types of membrane reactors with the scope of reducing the cost. Yu et al. from the Korea Research Institute of Chemical Technology used Pt-impregnated Knudsen membranes to carry out simultaneously the ESR reaction and WGS [START_REF] Yu | Ethanol steam reforming in a membrane reactor with Pt-impregnated Knudsen membranes[END_REF]. The ethanol reforming-membrane reactor showed ethanol conversion improvement up to ca. 15% in comparison with a conventional reactor, with an improvement of hydrogen yield up to 10.5%. A similar experiment was performed with a CMR loaded with Pt/TiO 2 catalyst, which showed hydrogen recovery values of 78-87% in the temperature range 573-873 K [START_REF] Yu | Study on a catalytic membrane reactor for hydrogen production from ethanol steam reforming[END_REF]. Oyama's group at the Virginia Polytechnic Institute & State University, US, employed silica-alumina composite membranes with moderate hydrogen permeance as well as Pd and Pd-Cu (β m) composite membranes over Al 2 O 3 for ESR over Na-Co/ZnO catalyst [START_REF] Lim | Reaction of primary and secondary products in a membrane reactor: Studies of ethanol steam reforming with a silica-alumina composite membrane[END_REF][START_REF] Lim | Studies of the effect of pressure and hydrogen permeance on the ethanol steam reforming reaction with palladium-and silica-based membranes[END_REF][START_REF] Yun | Experimental and kinetic studies of the ethanol steam reforming reaction equipped with ultrathin Pd and Pd-Cu membranes for improved conversion and hydrogen yield[END_REF]. Ethanol conversion and hydrogen production enhancement were measured to be ca. 20% in the Pd-Cu CMR. High-effective hydrogen production from ethanol and water was reported in a tubular dense mixedconducting oxygen permeable membrane reactor, in which the water splitting takes place at the tube side of the membrane and the oxidative steam reforming of ethanol occurs at the shell side simultaneously [START_REF] Zhu | Toward highly-effective and sustainable hydrogen production: bio-ethanol oxidative steam reforming coupled with water splitting in a thin tubular membrane reactor[END_REF]. Yttria stabilized zirconia hollow fibers were used as substrate for the deposition of a Pd-Ag membrane on the outer shell of the hollow fiber and for deposition of NiO/MgO-CeO 2 catalyst inside the hollow fiber to develop a catalytic hollow fiber membrane reactor for conducting ESR [START_REF] Rahman | On-board H2 generation by a catalytic hollow fibre microreactor for portable device applications[END_REF][START_REF] Rahman | Development of a catalytic hollow fibre membrane microreactor as a microreformer unit for automotive application[END_REF]. At 583 K, the hydrogen produced in the catalytic hollow fiber membrane reactor was twofold higher than that in the catalytic hollow fiber reactor without Pd-Ag membrane.

Effect of co-existence of ESR products on the permeation behavior of Pd-Ag membrane

As stated before, the influence of reforming products such as CO, CO 2 , H 2 O, and CH 4 on the hydrogen permeability of the Pd-based membranes (hydrogen inhibition) is considered as a drawback. Several works have been devoted to the investigation of the permeation behavior of the Pd-based membranes in the WGS and MSR environment [START_REF] Amandusson | The effect of CO and O2 on hydrogen permeation through a palladium membrane[END_REF][START_REF] Catalano | Influence of the gas phase resistance on hydrogen flux through thin palladium-silver membranes[END_REF][START_REF] Coroneo | Modelling the effect of operating conditions on hydrodynamics and mass transfer in a Pd-Ag membrane module for H2 purification[END_REF][START_REF] Gallucci | The effect of mixture gas on hydrogen permeation through a palladium membrane: Experimental study and theoretical approach[END_REF][START_REF] Israni | Methanol steam reforming in single-fiber packed bed Pd-Ag membrane reactor: Experiments and modeling[END_REF][START_REF] Li | The effect of carbon monoxide and steam on the hydrogen permeability of a Pd/stainless steel membrane[END_REF][START_REF] Nekhamkina | Effective approximations for concentrationpolarization in Pd-membrane separators[END_REF][START_REF] Peters T A | High pressure performance of thin Pd-23%Ag/stainless steel composite membranes in water gas shift gas mixtures; influence of dilution, mass transfer and surface effects on the hydrogen flux[END_REF][START_REF] Unemoto | The effect of co-existing gases from the process of steam reforming reaction on hydrogen permeability of palladium alloy membrane at high temperatures[END_REF]. Catalano et al. [START_REF] Catalano | Influence of the gas phase resistance on hydrogen flux through thin palladium-silver membranes[END_REF] stated that CH 4 and N 2 acted as inert gases in terms of hydrogen inhibition, although in their permeation tests the hydrogen flux drastically decreased when a mixture of hydrogen and N 2 and/or CH 4 was used. This was ascribed to the hydrogen partial pressure decrease in the case of binary or ternary mixtures. Several permeation tests were performed by Gallucci et al. [START_REF] Gallucci | The effect of mixture gas on hydrogen permeation through a palladium membrane: Experimental study and theoretical approach[END_REF] at 523-723 K and 2-3.5 bar using a palladium membrane to evaluate the influence of the presence of gases such as CO, N 2 , CO 2 , and Ar. The remarkable and quick influence of CO on the hydrogen permeation was reported, specially at low concentrations of CO. In the case of the rest of the gases, only the dilution effect was seen, by which the hydrogen flux was decreased but only due to the lower partial pressure of hydrogen (no membrane surface effect) [START_REF] Gallucci | The effect of mixture gas on hydrogen permeation through a palladium membrane: Experimental study and theoretical approach[END_REF]. The same results was reported by Li et al. [START_REF] Li | The effect of carbon monoxide and steam on the hydrogen permeability of a Pd/stainless steel membrane[END_REF] where a notable decrease in hydrogen permeation was observed at low concentrations of CO. Addition of steam to low concentrations of CO showed a stronger hydrogen inhibition.

Several authors mentioned the evident negative effect of presence of CO on the hydrogen permeance behavior of Pd-based membranes due to the covalent interaction of carbon monoxide with the membrane surface [START_REF] Israni | Methanol steam reforming in single-fiber packed bed Pd-Ag membrane reactor: Experiments and modeling[END_REF][START_REF] Catalano | Hydrogen permeation in palladiumbased membranes in the presence of carbon monoxide[END_REF][START_REF] Barbieri | A novel model equation for the permeation of hydrogen in mixture with carbon monoxide through Pd-Ag membranes[END_REF]. Peters et al. [START_REF] Peters T A | High pressure performance of thin Pd-23%Ag/stainless steel composite membranes in water gas shift gas mixtures; influence of dilution, mass transfer and surface effects on the hydrogen flux[END_REF] studied a Pd-Ag membrane reactor in WGS reaction environment at 673 K and reported the significant effect of CO on the hydrogen permeation rate through the membrane. Amandusson et al. [START_REF] Amandusson | The effect of CO and O2 on hydrogen permeation through a palladium membrane[END_REF] tested a β5 m thick Pd membrane in presence of hydrogen, CO, and O 2 at low temperatures of 373-523 K. They found out that both CO and O 2 influenced the membrane permeation performance. The available hydrogen dissociation sites were blocked by CO, and O 2 reacted with the present hydrogen to form water. The authors reported that at lower temperatures, CO dominated the effect of the formed water and blocked the permeation sites of the Pd membrane surface, resulting in less hydrogen permeation rate. Coroneo et al. [START_REF] Coroneo | Modelling the effect of operating conditions on hydrodynamics and mass transfer in a Pd-Ag membrane module for H2 purification[END_REF] developed a CFD model to compute the permeation rate of hydrogen through a Pd-Ag membrane in binary and ternary gas mixtures at 647-773 K and P=1-7 bar. The separation performance of the membrane was evaluated in presence of N 2 , CO, and CH 4 and at constant temperature and partial pressure of hydrogen. The negative effect of CO on the permeation flow rate of hydrogen was seen.

Mejdell et al. [START_REF] Mejdell | The effect of heat treatment in air on CO inhibition of a ∼γ m Pd-Ag (23wt.%) membrane[END_REF] investigated CO inhibition at 573-623 K on a very thin Pd-Ag membrane (γ m) before and after an air treatment done at 57γ K. The hydrogen flux declined drastically by 60% when the membrane was exposed to 1 mol% CO before the air treatment, while after the air treatment the same effect (CO inhibition) was 15 %. The enhanced performance of the membrane after the air treatment was attributed to the changes in CO and H 2 heats of adsorption due to the surface segregation of Pd or Ag [START_REF] Mejdell | The effect of heat treatment in air on CO inhibition of a ∼γ m Pd-Ag (23wt.%) membrane[END_REF].

As presented in the literature, CO has the dominant effect on the hydrogen permeation decrease by poisoning the permeation sites of the Pd metallic membranes.

Sanz et al. [START_REF] Sanz | H2 production via water gas shift in a composite Pd membrane reactor prepared by the pore-plating method[END_REF] carried out permeation tests using a Pd membrane reactor at 623-723 K and trans membrane pressure difference of 2.5 bar. The tests were done in the atmosphere of pure H 2 /N 2 , and mixed with CO or/and CO 2 . In all cases, a decrease in the hydrogen permeance was observed. However, the selectivity towards hydrogen was kept in all the cases. The results of the permeation tests at 623 K done by Sanz et al. [START_REF] Sanz | H2 production via water gas shift in a composite Pd membrane reactor prepared by the pore-plating method[END_REF] are given in Fig. 2.3. P H2ret and P H2perm represent the partial pressure of hydrogen at the retentate side and the permeate side, respectively.

got stable at a lower rate compared to the pure hydrogen stream. The authors claimed that in the WGS environment, the hydrogen inhibition effect of CO (the product of the reaction of H 2 O and CO 2 ) is not noticeable in front of the significant effect of carbon deposition on the surface of the membrane [START_REF] Gielens | Influence of steam and carbon dioxide on the hydrogen flux through thin Pd/Ag and Pd membranes[END_REF].

Hou et al. [START_REF] Hou | The effect of external mass transfer, competitive adsorption and coking on hydrogen permeation through thin Pd/Ag membranes[END_REF] reported that the hydrogen inhibition effect of CO, CO 2 , and H 2 O in the case of a Pd-Ag membrane could be classified as H 2 O>CO>>CO 2 . The order of influence corresponded to the competitive adsorption capability of the gases on the Pd-Ag membrane surface. At temperatures higher than 673 K, the hydrogen inhibition was not noticeable. In the study by Unemoto et al. [START_REF] Unemoto | The effect of co-existing gases from the process of steam reforming reaction on hydrogen permeability of palladium alloy membrane at high temperatures[END_REF] the comparison between CO, CO 2 , and H 2 O showed that at T<600 K, CO had the strongest influence on the hydrogen permeability of the Pd membrane. They suggested that at T>873 K, the effect of co-existence of other species for a membrane with a thickness higher than 10 m was negligible. On the contrary, Patrascu and Sheintuch [START_REF] Patrascu | On-site pure hydrogen production by methane steam reforming in high flux membrane reactor: Experimental validation, model predictions and membrane inhibition[END_REF] concluded that the effect of very small amount of CO on hydrogen permeation inhibition could be notable even in presence of H 2 O, and especially in the steam reforming environment.

Overall, the published open literature offers no robust model/analysis on the effect of the different species on the performance of the membrane in the real atmosphere of methane steam reforming and water gas shift reactions. The competitive adsorption of CO and H 2 O on the surface of the metallic Pd membrane, the effect of reverse reactions of WGS and MSR, and the effect of operating at high pressure and temperature do not give a consistent idea of the hydrogen inhibition of the ESR products in the real atmosphere of the steam reforming of ethanol over noble metals. Moreover, there is not a consistent conclusion on the effect of H 2 O on the permeability or active surface area of the membrane. It is not totally agreed if CO 2 and CH 4 are considered as inert gases, as the decomposition of these two species, or the reactions with water via WGS and MSR, respectively, can lead to a more complicated situation regarding the influence on the hydrogen permeation. According to the review given by Cornaglia et al. [START_REF] Cornaglia | Recent advances in catalysts, palladium alloys and high temperature WGS membrane reactors[END_REF], it can be understood that the hydrogen inhibition phenomenon caused by the ESR products; specially in presence of H 2 O, is a very complicated issue. Specific considerations in each case in terms of the properties of the membrane, operating conditions, and the composition of the fuel fed into the reactor (S/C ratio) may be a relevant solution.

In this chapter an introduction to the membrane reactors and a literature review on the ethanol steam reforming in the membrane reactors over noble metal based catalyst were given. While many interesting studies have been performed recently on the application of membrane reactors in pure hydrogen production, this area requires further research and development to introduce efficient, robust, and applicable pure hydrogen-producing systems. The next chapter is devoted to the experimental investigation of the ethanol steam reforming reaction over Pd-Rh/CeO 2 in a Pd-Ag membrane reactor at different operating conditions to study the effect of catalyst reactivity, hydrogen selectivity, fuel conversion, and pure hydrogen permeation rate through the Pd-Ag membrane.

Chapter 3 Experimental results of ethanol steam reforming over Pd-Rh/CeO 2 in the Pd-Ag membrane reactor

Abstract

A comprehensive introduction and literature review on hydrogen, membrane reactors, and common catalysts for hydrogen production via catalytic steam reforming of ethanol was given in previous chapter. As the first step, ethanol steam reforming (ESR) experiments over Pd-Rh/CeO 2 catalyst were performed in a catalytic membrane reactor (CMR) at different temperatures and pressures, using a mixture of ethanol and water at various S/C ratios. The reactor was filled by the catalyst so that the membranes were fully covered to achieve the maximum amount of hydrogen production and permeation. Hydrogen yield of 0.6 at 12 bar and 923 K was reached at S/C=3 and hydrogen recovery was 92%. 0.9 L N hydrogen/ml EtOH equal to 3.5 mole of pure hydrogen per mole of inlet EtOH was obtained at 12 bar. The prototype was established so that pure hydrogen was obtained as it was generated, thanks to the high selectivity of the Pd-Ag membranes towards hydrogen.

Hydrogen was obtained at the purity of 99.999%, suitable for direct online use for low temperature fuel cells [START_REF] Koch | Ethanol catalytic membrane reformer for direct PEM FC feeding[END_REF]. The experimental results were compared to a staged membrane reactor (SMR) [START_REF] López | Hydrogen production from ethanol over Pd-Rh/CeO2 with a metallic membrane reactor[END_REF][START_REF] Koch | Ethanol catalytic membrane reformer for direct PEM FC feeding[END_REF] configuration in which the same experimental conditions were applied and the same catalyst was used.

Differently, at temperatures higher than 573 K, the dissociation rate is limited by the diffusion of the hydrogen atoms through the palladium layer, where the hydrogen flux follows the Sievert's law. Based on Fig. 3.8 and considering a membrane area of 30.4 cm 2 , the hydrogen permeation rate at λβγ K (650˚C) and different pressures is obtained, as presented in Table 3 The maximum and minimum pressures at which the permeation test in pure hydrogen environment was performed are 20 and 3.4 bar [START_REF] Papadias | An analytical and experimental investigation of high-pressure catalytic steam reforming of ethanol in a hydrogen selective membrane reactor[END_REF]. The maximum pressure at which the ESR reaction in the CMR were performed in this work is 12 bar, and finally the maximum partial pressure of hydrogen reached in this work was around 7.2 bar (in the ESR environment in the CMR). The values presented in Table 3.2 are the highest hydrogen permeation rates possible through the studied membrane at mentioned temperature and pressure.

Comparison between two reactor configurations

In addition to the evaluation of the CMR configuration, a comparison to the previously studied reactor configuration named as the Stage Membrane Reactor (SMR) [START_REF] López | Hydrogen production from ethanol over Pd-Rh/CeO2 with a metallic membrane reactor[END_REF][START_REF] Koch | Ethanol catalytic membrane reformer for direct PEM FC feeding[END_REF] is given.

In the CMR configuration, the membrane was fully covered by the catalyst, therefore, the ESR reaction and the hydrogen permeation occurred simultaneously in the CMR. In the case of the SMR, the catalytic zone was separated from the permeation zone, which was the membrane, so that firstly the ESR reaction were complete and came to equilibrium, then the gaseous products passed along the membrane letting the hydrogen to permeate through the membrane. The two different configurations are shown in Fig. 3.9. P H2 represents the partial pressure of hydrogen inside the reactor at the membrane wall. The partial pressure of hydrogen at the permeate side was one bar (no sweep gas used).

In the CMR, the dynamics is different. Up until the retentate gas exits the reactor, the reforming reactions are in progress considering the shift effect. In the CMR, the higher the pressure, the higher the hydrogen permeation rate, and the higher the production rate of hydrogen. In fact, the thermodynamic equilibrium limits are conquered owing to the shift effect and the brilliant activity of the catalyst to convert largely the fuel to CO 2 and H 2 . It is assumed that the partial pressure of hydrogen is kept high along the membrane because of simultaneous and continuous hydrogen production.

The selectivity of the retentate gas (outlet gas) at the same experimental conditions also proves the superior performance of the CMR. The molar flow rates of CO and CH 4 as a function of pressure for both configurations are shown in Fig. 3.24.

Conclusion

ESR experiments over Pd-Rh/CeO 2 were performed in a CMR containing Pd-Ag separation membranes using ethanol and water mixtures at different S/C ratios. The effect of operating conditions, specially temperature and pressure, was discussed. The performance of the CMR was studied in terms of pure hydrogen production, hydrogen yield, and hydrogen recovery. The shift effect resulted in very high hydrogen production and the equilibrium limitations at high pressure were conquered. The influence of the coexistence of species such as CO and CH 4 was shown based on the experimental results of the permeation tests in the atmosphere of pure hydrogen and ethanol steam reforming. A comparison between the CMR and the SMR clearly revealed the superior performance of the CMR. Many observed features in the case of SMR were not seen in the CMR so that pure hydrogen production rate and methane conversion were promoted at high pressure.

To conclude, some of the highlights of the experimental work can be listed as:

• Complete ethanol conversion at all the operating conditions.

• Hydrogen yield of 0.55 at 923 K, 12 bar, and S/C=3 in the CMR (0.2 in the SMR).

• Hydrogen recovery reaches up to 95% at 923 K, 12 bar, and S/C=3 in CMR (70% in SMR). • 0.9 L N Hydrogen/ml EtOH at 12 bar, and 923 K.

• More than 3 moles of pure hydrogen per mole of inlet EtOH.

• ESR reaction promotion beyond the equilibrium limitations at high pressure.

• Positive effect of pressure on the performance of the CMR despite unfavorable thermodynamic conditions at high pressure.

The next chapter presents the results of the thermodynamic analysis of the reforming system mainly in terms of thermal and exergy efficiency based on the experimental results given in this chapter. Both CMR and SMR configurations are discussed and compared in detail as a complement to the experimental study, to provide a detailed overview of the discussed reforming system.

Chapter 4 Thermodynamic analysis of the CMR and SMR systems

Abstract

The experimental results of the ESR reaction in the CMR and the effect of various operating conditions on the hydrogen selectivity and pure hydrogen production rate were given in previous chapter. Aiming to understand the thermodynamic performance of the reformer, the energy and exergy analysis are given in this chapter. Thermodynamic evaluation is presented as a supplement to the comprehensive investigation of the performance. The experimental results presented in the previous chapter (selectivity of ESR products and the pure hydrogen production rate at all the presented operating conditions) were used to perform the exergy analysis. The exergy efficiency and exergy destruction of the CMR were evaluated based on inlet and outlet streams. The exergy efficiency of the reformer was evaluated considering both an insulated reactor (without heat loss), and a non-insulated reactor (with heat loss). The significant effect of pressure and temperature on the exergy efficiency was observed. Exergy efficiency up to 50% was reached in the case of an insulated reactor at 12 bar and 923 K. It was concluded that the highest rate of exergy destruction occurred via heat losses. The study showed that the exergy content of the retentate gas especially at lower pressures could provide the reactor with a notable fraction of its required heat at steady state conditions. This notably increased the exergy efficiency of the reformer. In the case of recovery of the retentate gas and insulation of the reactor, exergy efficiency placed between 70-90%. It was concluded that operating at the highest pressure, the lowest S/C ratio, and at 923 K gives the best exergy efficiency. Thermal efficiency of the reformer was between 60-90% for an insulated reactor and decreased to between 40-60% when the heat loss was considered. The exergy analysis proved superior performance of the CMR.

Exergy analysis

Regarding the thermodynamic investigations of the ESR process, the studies reported in the literature are mainly focused on energy efficiency (ratio of energy output to energy input), but not on exergy analysis. Exergy analysis is a powerful tool to investigate the imperfections of single components of an energy system to obtain a better understanding of the local irreversibility by means of a more detailed picture of the process.

The term exergy is defined as the maximum work that can be obtained theoretically from a system interacting with the source environment to equilibrium [START_REF] Adrian Bejan | Thermal Design and optimization[END_REF]. Exergy differs from energy in the way that energy is conserved, but exergy can be dissipated. Exergy is defined based on both the first and second laws of thermodynamics stating that it is not possible to fully utilize the thermal energy [START_REF] Sato | Chemical Energy and Exergy[END_REF]. In other words, exergy is the ability of available energy to convert into other forms of energy.

Taking into account the second law of thermodynamics, exergy is derived from the entropy, free energy (Helmholtz energy), and free enthalpy (Gibbs free enthalpy). Therefore, exergy is a function of the thermodynamic state of the substance and the reference environment [START_REF] Sato | Chemical Energy and Exergy[END_REF]. In the light of exergy definition, it can be understood that the main difference between energy (thermal) efficiency and exergy efficiency lies in the consideration of the thermodynamic state of every single component, which results in an exact understanding of the available amount of work, together with the unavoidable irreversibility during a process. Exergy analysis makes it possible to study the effect of thermodynamic factors on the performance of an energy system to decide on the most favorable operational conditions in terms of process efficiency and energy usage [START_REF] Khila | A comparative study on energetic and exergetic assessment of hydrogen production from bioethanol via steam reforming, partial oxidation and auto-thermal reforming processes[END_REF][START_REF] Casas-Ledón | Thermodynamic Analysis of the Hydrogen Production from Ethanol: First and Second Laws Approaches[END_REF].

According to the open literature, there are a few reported studies on exergy efficiency evaluation for hydrogen production. Kalinci et al. [START_REF] Kalinci | Efficiency assessment of an integrated gasifier/boiler system for hydrogen production with different biomass types[END_REF] studied the production of hydrogen via a gasification-boiler system based on experimental data taken from the literature using different types of biomass. They found the maximum exergy efficiency to be about 12%. An exergy analysis of the biological hydrogen production from biomass was done by Modarresi et al. [START_REF] Modarresi | Application of exergy balances for evaluation of process configurations for biological hydrogen production[END_REF] based on experimental results. They reported exergy efficiencies of 36-45%, depending on the process configuration. For reforming processes, Simpson et al. [START_REF] Simpson | Exergy analysis of hydrogen production via steam methane reforming[END_REF] modelled the methane steam reforming process and both irreversible chemical reactions and heat losses were identified as the main source of exergy destruction, whereas exhaust gases contained large amounts of chemical exergy. Casas-Ledón et al. [START_REF] Casas-Ledón | Thermodynamic Analysis of the Hydrogen Production from Ethanol: First and Second Laws Approaches[END_REF] studied hydrogen production from ESR based on the first and second laws of thermodynamics. They evaluated the exergy efficiency of the system experimentally at different operational conditions (pressure, temperature, and S/C ratio) considering the unused and destructed exergy during the ESR process. They concluded that the exergy efficiency of the ESR system was a function of temperature and S/C ratio, while no effect of pressure on exergy efficiency was observed. A comprehensive exergy analysis of the different types of ethanol reforming processes (ESR, POX and ATR) based on a model in Aspen Plus was performed by Khila et al. [START_REF] Khila | A comparative study on energetic and exergetic assessment of hydrogen production from bioethanol via steam reforming, partial oxidation and auto-thermal reforming processes[END_REF]. The same formulation as Casas-Ledón et al. [START_REF] Casas-Ledón | Thermodynamic Analysis of the Hydrogen Production from Ethanol: First and Second Laws Approaches[END_REF] was used by Khila et al. and they used Aspen Plus software to calculate the exergy of the inlet and outlet streams at selected operational conditions, according to hydrogen production per mole of inlet ethanol. An exergy efficiency of 70% was claimed for the ESR process, considering total hydrogen production via ESR as the main product. In another study, Tippawan et al [START_REF] Tippawan | Energy and exergy analysis of an ethanol reforming process for solid oxide fuel cell applications[END_REF] employed the first and second law of thermodynamics to evaluate energy and exergy performance of a modelled ethanol reforming system in connection with a solid oxide fuel cell (SOFC) with a similar formulation as Casas-Ledón et al. [START_REF] Casas-Ledón | Thermodynamic Analysis of the Hydrogen Production from Ethanol: First and Second Laws Approaches[END_REF] and Khila et al [START_REF] Khila | A comparative study on energetic and exergetic assessment of hydrogen production from bioethanol via steam reforming, partial oxidation and auto-thermal reforming processes[END_REF]. They studied ESR, partial oxidation (POX), and autothermal reforming (ATR) processes as the reforming sections for hydrogen production, and the best efficiency of the system (reforming+SOFC) was stated equal to 60% when ESR was used as the reformer unit.

In this chapter, the application of exergy analysis in the evaluation of the ethanol steam reforming (ESR) process in the CMR and SMR configurations is presented. The molar flow rates of the pure hydrogen stream and all the ESR products (CO, H 2 , CH 4 , and CO 2 in the retentate stream) obtained via experiments were used for calculation of the exergy rate of different streams. Exergy evaluation was used to study the effect of the operating conditions on the performance of the system and to understand the thermodynamic losses, to decide on the best operating conditions in terms of both pure hydrogen production and the exergy efficiency.

Materials and methods

The traditional method of process performance evaluation is based on the first law of thermodynamics, which considers the lower heating value (LHV) of the inlet and outlet streams, and the amount of work or heat provided to run the process. Thermal efficiency of the reformer system is defined as [START_REF] Hedayati | Exergetic study of catalytic steam reforming of bio-ethanol over Pd-Rh/CeO2 with hydrogen purification in a membrane reactor[END_REF][START_REF] Khila | A comparative study on energetic and exergetic assessment of hydrogen production from bioethanol via steam reforming, partial oxidation and auto-thermal reforming processes[END_REF]:

ℎ = ̇ × + ṁ ℎ × ℎ ṁ × + ẇ + ̇ 4.1 W ̇
is the work of the pump. The work of the pump is neglected in this work. Q ̇ represents the rate of heat losses and the required energy to run the ESR in the CMR. Heat losses account for the heat released to the environment through the reactor wall, products cooling down, and water condensation. All the products (the retentate gas stream and condensed water) are cooled down by transferring heat to the environment via the pipes and the condenser wall.

Reactor wall temperature (in contact with air) was measured by means of a K-type thermocouple to calculate the heat loss at different operating temperatures (873 and 923 K). The heating band is the heat source for the evaporation of the fuel and heating up to pressure inside the reactor and outside the membrane (Sieverts' law). Nevertheless, this phenomenon is considered as energy neutral because the pressure needed is supplied by the pump.

Exergy efficiency is calculated based on exergy destruction and unused exergy, as a function of the exergy rate of the inlet and outlet streams. This formulation has been repetitively used by different researchers [START_REF] Hedayati | Exergetic study of catalytic steam reforming of bio-ethanol over Pd-Rh/CeO2 with hydrogen purification in a membrane reactor[END_REF][START_REF] Khila | A comparative study on energetic and exergetic assessment of hydrogen production from bioethanol via steam reforming, partial oxidation and auto-thermal reforming processes[END_REF][START_REF] Casas-Ledón | Thermodynamic Analysis of the Hydrogen Production from Ethanol: First and Second Laws Approaches[END_REF][START_REF] Cohce | Energy and exergy analyses of a biomass-based hydrogen production system[END_REF]. Exergy destruction rate is defined as:

= - 4.3
Where EX in and EX out are the exergy rates of the inlet and outlet streams. Therefore, EX in represents the exergy rate of inlet fuel, i.e. ethanol (with its specific exergy ex fuel ), plus the required exergy rate for the ESR process including heat losses (EX ̇ = W ̇ ), and EX out denotes the exergy rate of pure hydrogen stream (ex H2.perm ), plus the retentate gases exiting the reactor (ex retentate ):

= + ̇ 4.4 = . . + 4.5
The condensed water is considered to have zero exergy value. The rate of unused exergy is calculated as:

= - 4.6
Where EX retentate is the exergy rate of the retentate gas. The exergy of the pure hydrogen stream is considered as useful part of exergy. Accordingly, the fraction of hydrogen in the retentate gas (not permeated) is not taken into account as the main product. Finally, the exergy efficiency of the ESR process is given by eq. 4.7:

= 1 -

4.7

The specific exergy content of each species (ex) in each stream consists of physical exergy (ex physical ), chemical exergy (ex chemical ), and mixing exergy (ex mixing ):

= ℎ + ℎ + 4.8
Physical exergy (ex physical ) is the maximum obtainable work produced when a stream is brought from the actual conditions (T, P) to the reference conditions (P 0 , T 0 ) and is defined as [START_REF] Freni | Ethanol steam reforming in a molten carbonate fuel cell: a thermodynamic approach[END_REF][START_REF] Verne | Twenty thousand leagues under the sea[END_REF] :

h and s are the specific enthalpy and specific entropy of the substance at actual (reaction) conditions, and h 0 and s 0 are the specific enthalpy and specific entropy of the substance at reference conditions, respectively. The dependency of the physical exergy on enthalpy and entropy highlights two features. First, exergy is a function of the state of the matter, and second, each matter is considered independently in a stream. Both features result in a more precise idea on the performance of a thermal system. Chemical exergy originates from the difference between the chemical potentials when a substance is changed at reference conditions to the chemical equilibrium state with the concentrations of components. In this work, the chemical exergy of each specie was calculated using the standard chemical exergy table given by Bejan model II [START_REF] Adrian Bejan | Thermal Design and optimization[END_REF]. Chemical exergy occasionally is reported as a sum of two terms, i.e. the standard chemical exergy plus a logarithmic term as a function of the fraction of each substance in a mixture [START_REF] Casas-Ledón | Thermodynamic Analysis of the Hydrogen Production from Ethanol: First and Second Laws Approaches[END_REF][START_REF] Tippawan | Energy and exergy analysis of an ethanol reforming process for solid oxide fuel cell applications[END_REF]:

ℎ = + 4.12
Where x i is the fraction of any species in the mixture of gases, i is the standard chemical exergy of the species, and R is the universal gas constant. The physical properties together with standard chemical exergy for the chemical species are presented in Table 4 The second term in eq. 4.12 (RT x Lnx ), as is always negative, can be ascribed to the exergy of mixing. Exergy of mixing is the entropy generated when pure substances are mixed and is given by eq. 4.13 [START_REF] Sato | Chemical Energy and Exergy[END_REF]:

= 4.13
However, the value of mixing exergy is normally negligible in front of standard chemical exergy [START_REF] Hinderink | Exergy analysis with a flowsheeting simulator-I. Theory; calculating exergies of material streams[END_REF]. A comprehensive discussion on various types of exergy calculation is given by Sato [START_REF] Sato | Chemical Energy and Exergy[END_REF] and Hinderink et al. [START_REF] Hinderink | Exergy analysis with a flowsheeting simulator-I. Theory; calculating exergies of material streams[END_REF]. Similar definitions have been reported in some studies, which are based on the entropy difference between the mixture of substances and the pure components (which exist in the mixture) individually [START_REF] Khila | A comparative study on energetic and exergetic assessment of hydrogen production from bioethanol via steam reforming, partial oxidation and auto-thermal reforming processes[END_REF][START_REF] Hinderink | Exergy analysis with a flowsheeting simulator-I. Theory; calculating exergies of material streams[END_REF]. In this work, all three types of exergy were considered for each species in the inlet and outlet streams. The molar flow rate of reactants and products obtained during the experimental work were used to calculate the exergy rate of each stream.

Results and Discussion

Thermodynamic analysis of the CMR

The optimization of the reforming system not only depends on the improvement of pure hydrogen permeation rate, but also on the thermodynamic efficiencies. Thermodynamic evaluation provides a better understanding of the effect of the operating conditions on the performance of the reformer. Pure hydrogen production gets higher naturally with the fuel flow rate, as discussed earlier. This is in agreement with the thermal efficiency, but not with exergy efficiency, due to the criteria based on which the thermal and exergy efficiencies are defined (eq. 4.1 and 4.7, respectively). Thermal and exergy efficiency at FF=50 and 100 SMR, but in the case of the CMR, two different behavior is seen. At FF=50, the exergy efficiency declines with pressure, which can be a result of higher exergy rate required for the ESR reaction at recovery of the retentate gas when the reactor in insulated. However, at FF=100, the hydrogen production rate is high enough to improve the exergy efficiency. It is seen that, generally, exergy efficiency of the SMR is higher than CMR, but it must be taken into acount that the exergy efficiency is taken into account together with the rate of pure hydrogen production, methane conversion, and hydrogen yield and recovery. Accordingly, at high pressure, the performance of the CMR is advantageous in terms of pure hydrogen production, low unused exergy rate, and high exergy efficiency.

Conclusion

Exergy analysis of the CMR and SMR configuration was performed in terms of exergy and thermal efficiency based on the experimental results at various operating conditions. It was concluded that exergy analysis is a relevant criteria for evaluating the reforming system. The advantages of exergy evaluation over thermal efficiency study was stated according to the explanation of the effects of operating conditions on both pure hydrogen production rate, and total efficiency of the CMR.

An exergy analysis was performed based on the experimental results aiming not only to evaluate the performance of the CMR system, but also to introduce the application of the exergy analysis in CMRs studies. Exergy analysis provided important information on the effect of operating conditions and thermodynamic losses, resulted in understanding of the best operating conditions. Both insulated and non-insulated reactor systems were evaluated in terms of exergy destruction, exergy efficiency, and thermal efficiency.

The effects of pressure and temperature were dominant and the study showed that the system reached around 50% exergy efficiency at 923 K and 12 bar in an insulated reactor. Unused exergy decreased with pressure since the MSR reaction is promoted at high pressures in CMR as more hydrogen is permeated due to the availability of the catalyst around the membrane and the shift effect. The exergy content of the retentate gas could compensate the energy requirements of the reactor to improve the exergy efficiency significantly. The highest exergy destruction occurred via heat losses and the retentate stream. In the case of recovery of the retentate gas and insulation of the reactor, exergy efficiency of the CMR placed between 60-90%. It was concluded that operating at the highest pressure, the lowest S/C ratio and at 923 K gives the best exergy efficiency for the CMR. Although operating at higher fuel flow rate is more beneficial regarding the production rate of pure hydrogen, the exergy efficiency is slightly lower.

The thermal efficiency of the CMR was also evaluated based on the LHV of ethanol and the products and was compared to the exergy efficiency. While thermal efficiency offered an ideal performance at high pressure and temperature, exergy efficiency disclosed inevitable irreversibilities even in the case of an insulated reactor.

The thermodynamic analysis of the SMR showed different results so that the exergy and thermal efficiency were lower compared to the CMR. This was attributed to lower hydrogen production, more methane flow rate in the retentate stream (lower methane conversion in the case of the SMR), and very large heat loss area (high heat loss rate) of the SMR. The exergy rate of the retentate stream in the case of the SMR was higher than the CMR due to the low methane conversion in the SMR.

At least 80% of the inlet exergy was lost by heat losses and the retentate stream in the SMR. The insulation of the SMR resulted in a significant increase in the thermal efficiency, but not exergy efficiency because of low methane conversion and hydrogen recovery (high rate of unused exergy).

Superior performance of the CMR was proved thermodynamically. Taking into account the pure hydrogen production rate and exergy efficiency at the same time, the CMR showed an overall enhanced performance. In the case of the recovery of the retentate gas in an insulated reactor, the exergy efficiency of the SMR was slightly higher than the CMR, however, much higher hydrogen yield and recovery was reached in the case of the CMR.

As the area of the membrane science and pure hydrogen production in membrane reactors is growing, exergy evaluation of the CMR systems -as the first essential step for system analysis via exergo-economical optimization -can open a new chapter in this science to approach larger scale applications.

The next chapter is devoted to the modeling and dynamic simulation of the reforming system specially at transient state where the operating pressure or fuel flow rate is adjusted. The dynamic simulation can be considered as modeling the response of the reforming system to a fuel cell -when fed online -in terms of pure hydrogen production rate modification.

130 Chapter 5

Modeling and dynamic simulation of the reforming system

Abstract

In chapters 3 and 4, the results of the experimental and thermodynamic analysis of the reformer were discussed. In this chapter, the simulation of the dynamics of hydrogen production (permeation) is presented as the last step to study the applicability of such a system in connection with a real end user, which can be a fuel cell. The simulation presented in this chapter is similar to the hydrogen flow rate adjustments needed to set the electrical load of a fuel cell, when fed online by the reformer. Ethanol steam reforming experiments were performed at 923 K, 6-10 bar, and fuel flow rates of 50 to 200 µl/min using a mixture of ethanol and distilled water with steam to carbon ratio of 3. Dynamic experiments were carried out to observe the behavior of the CMR in terms of the pure hydrogen production rate in the case of operating pressure or fuel flow rate adjustments. A static model for the catalytic zone was derived from the Arrhenius law as a function of operating conditions to simulate the total molar production rates of ESR products. The pure hydrogen production rate at steady state conditions was simulated by means of two static models named as model 1 and model 2. Model 1 represented the simulation based on the Sieverts' law, which is known as the physical definition of hydrogen permeation through the membrane via which the robustness of the simulation was proved. Model 2 represented a black box model as a function of total hydrogen production via ESR, and the reactor pressure. Finally, a dynamic model was proposed under the ideal gas law assumptions to simulate the dynamics of pure hydrogen production rate in transient state at isothermal conditions. Both pressure and fuel flow rate change steps simulations fitted the experimental values very well. However, the simulation of the dynamics of the fuel flow rate change was more essential, as the system responds much faster to such an adjustment. The dynamic simulation proved a successful essential step needed to simulate an "Ethanol to Electricity" system.

Dynamic simulation of the CMRs

A large piece of work on ESR in CMRs using different catalysts and reactor configurations can be found in the literature [START_REF] Llorca | Alcohols and Bio-alcohols Steam and Autothermal Reforming in a Membrane Reactor[END_REF]. Generally, ESR in membrane reactors has been simulated based on the experimental results, aiming to mathematical modeling and kinetic studies. The reported models have been derived from conventional power-law, at different operating conditions (temperature, pressure, and S/C ratio) mainly to model the reaction rates, ethanol conversion, and the selectivity of reforming products [START_REF] Palma | CeO2-supported Pt/Ni catalyst for the renewable and clean H2 production via ethanol steam reforming[END_REF][START_REF] Tosti | Low temperature ethanol steam reforming in a Pd-Ag membrane reactorPart 1: Ru-based catalyst[END_REF][START_REF] Papadias | An analytical and experimental investigation of high-pressure catalytic steam reforming of ethanol in a hydrogen selective membrane reactor[END_REF][START_REF] Bej | Production of hydrogen by steam reforming of ethanol over alumina supported nano-NiO/SiO2 catalyst[END_REF][START_REF] Wang | From mechanistic to kinetic analyses of ethanol steam reforming over Ir/CeO2 catalyst[END_REF][START_REF] López | Ethanol steam reforming for hydrogen generation over structured catalysts[END_REF][START_REF] Gallucci | Co-current and counter-current configurations for ethanol steam reforming in a dense Pd-Ag membrane reactor[END_REF].

Some modeling studies followed by experimental validations have been devoted to the design and optimization of the water gas shift (WGS) reaction in CMRs containing Pd composite membranes (mostly Pd-Ag). These studies aimed to produce pure hydrogen and monitoring the effect of temperature, catalyst loading, steam to carbon ratio S/C, and permeate side pressure on the performance of the CMRs [START_REF] Hla | Modelling and experimental studies of a watergas shift catalytic membrane reactor[END_REF][START_REF] Basile | Water gas shift reaction in membrane reactors: Theoretical investigation by artificial neural networks model and experimental validation[END_REF][START_REF] Sanz | Hydrogen production in a Pore-Plated Pd-membrane reactor: Experimental analysis and model validation for the Water Gas Shift reaction[END_REF]. Dong et al. [START_REF] Dong | Tubular dual-layer MFI zeolite membrane reactor for hydrogen production via the WGS reaction: Experimental and modeling studies[END_REF] studied hydrogen production via WGSR in a membrane reactor with a tubular ZSM-5/silicalite bilayer membrane. They optimized the process by means of modeling so that they could reach high CO conversion (>95%) and H 2 recovery (>90%) for WGS in the zeolite membrane reactor. A three-dimensional numerical model was developed by Chein et al. [START_REF] Chein | Three-dimensional numerical modeling on high pressure membrane reactors for high temperature water-gas shift reaction[END_REF] to simulate hydrogen production from coal-derived syngas via the WGS reaction in membrane reactors at 1173 K. High temperature was chosen to simulate the typical temperature of the syngas at the exit of a gasifier. The negative effect of CO 2 content on the hydrogen permeation rate was observed. In addition to WGS reaction, methane steam reforming (MSR) reaction have been investigated by performing experiments and development of models [START_REF] Patrascu | On-site pure hydrogen production by methane steam reforming in high flux membrane reactor: Experimental validation, model predictions and membrane inhibition[END_REF][START_REF] Marcoberardino | Fixed bed membrane reactor for hydrogen production from steam methane reforming: Experimental and modeling approach[END_REF][START_REF] Iulianelli | Model biogas steam reforming in a thin Pd-supported membrane reactor to generate clean hydrogen for fuel cells[END_REF]. In this regard, the adverse effects of the presence of different species such as CH 4 , CO, and H 2 O on the permeation rate of membrane have been studied. WGS reaction and MSR are the two determining reactions when ESR is performed over Pd-Rh/CeO 2 catalyst used also in the present work.

If a fuel cell is fed online by a pure hydrogen generating system (reformer), the dynamics of the pure hydrogen supply must be fitted to the load variations (dynamic behavior) of the fuel cell. Considering the dynamic energy demand of an end user -for example a building -a reformer must be able to realize and track the dynamic electrical output of the fuel cell in charge of electricity supply of the end user. Adjustment of the flow rate of pure hydrogen provided by a reformer is a crucial phase to respond promptly and aptly to the electrical load modifications of a fuel cell, aiming to optimize the whole system (reformer + fuel cell) performance. Although some studies are reported in the literature regarding the dynamic performance of the fuel cells, works devoted to the investigations of the dynamic performance of the online fuel reformers -corresponding to the load variation of the fuel cells -are not sufficiently reported [START_REF] Zhang | Dynamic performance of a hightemperature PEM fuel cell -An experimental study[END_REF].

Garcia et al. [START_REF] García | Dynamic modeling of a three-stage lowtemperature ethanol reformer for fuel cell application[END_REF] developed a dynamic model for a three module reformer made up of ethanol dehydrogenation, acetaldehyde steam reforming, and water gas shift units for feeding hydrogen to a fuel cell. They simulated the dynamic response of the reforming unit in terms of the selectivity of the products of the ESR reaction rate to the changes in concentration of the feed (ethanol + water). The same authors in another study [START_REF] García | Dynamic modeling and controllability analysis of an ethanol reformer for fuel cell application[END_REF] focused on the controllability and the dynamic simulation of the same system as they reported in [START_REF] García | Dynamic modeling of a three-stage lowtemperature ethanol reformer for fuel cell application[END_REF] by acting on the feed concentration at isothermal conditions. A dynamic numerical model for the methane fuel processor of a PEMFC was developed by Funke et al. [START_REF] Funke | A dynamic model of the fuel processor for a residential PEM fuel cell energy system[END_REF] aiming at optimizing the reaction conditions and heat integration especially during start up, shut down, and load change. The effect of two constructions (the reactor and the evaporator with and without thermal coupling) on the temperature profile, reaction rates, and methane conversion was investigated. Hydrogen yield was higher when the reactor and the evaporator were not thermally coupled. John and Schroer [START_REF] Jahn | Dynamic simulation model of a steam reformer for a residential fuel cell power plant[END_REF] presented a dynamic model of a methane steam reformer for a residential fuel cell system. The dynamic model covered the full operating range including the startup and shut down, and described the dynamics of the hydrogen yield and thermal behavior of the reformer when the flow rate of water or natural gas changed. The thermal system was affected by increasing the flow rate of the water. Higher hydrogen yield and lower methane concentration at the outlet were reported at higher temperature, i.e. lower concentration of inlet water. A dynamic model for an interconnected reformer and PEMFC stack was developed by Stamps and Gatzke [START_REF] Stamps | Dynamic modeling of a methanol reformer-PEMFC stack system for analysis and design[END_REF] with emphasis on the influence of various design and operating parameters on system performance. Operating at higher temperature resulted in higher system performance.

A dynamic modeling study of a catalytic steam reformer by Kvamsdal et al. [START_REF] Kvamsdal | Dynamic simulation and optimization of a catalytic steam reformer[END_REF] showed that the steam or gas (CO, CO 2 , H 2 , and CH 4 ) supply interruption affects the reactor wall temperature, which can directly lead to material failure or coke formation. Lin et al. [START_REF] Lin | Dynamic modeling and control structure design of an experimental fuel processor[END_REF] modeled the dynamics of an experimental multi-stages methane reformer in charge of providing hydrogen to a PEMFC to design a control system to provide the responsiveness of the fuel reformer to the alterations in the hydrogen demand. The response of the fuel reformer to changes in the process variables such as CH 4 feed flow rate, H 2 O/CH 4 feed ratio, O 2 /CH 4 feed ratio and the reformer inlet temperature was studied. Tsourapas et al. [START_REF] Tsourapas | Modeling and dynamics of an autothermal JP5 fuel reformer for marine fuel cell applications[END_REF] presented a dynamic model based on thermodynamics and energy balance for a JP5 fuel reformer in connection with a membrane separator (SEP) and a PEM fuel cell to investigate the effects of the operating set point of SEP on the overall system efficiency. They concluded that the open loop response of the system is shown to be satisfactory in terms of the response time and hydrogen production. It was shown that there is a trade-off between the SEP efficiency and the overall efficiency of the system.

In another work by Koch et al. [START_REF] Koch | Ethanol catalytic membrane reformer for direct PEM FC feeding[END_REF], a dynamic model of an ethanol steam reformer (as the fuel reformer for pure hydrogen production to feed a PEMFC) was developed to implement an adaptive and predictive control. The static behavior of the reformer system was described by means of several maps developed in Matlab. Further, the dynamics of the fuel reformer (SMR) in connection with a PEMFC by acting upon reactor pressure and feed flow rate (ethanol + water) was studied. They proposed an efficient controller that reduced the response time of the reformer by a factor of 7 down to 8 s in terms of following the dynamic of a fuel cell load by acting simultaneously on the fuel flow rate and pressure. However, such advanced controllers require internal models and simulations for further development.

The purpose of this chapter is to present a simple approach mainly based on physical laws (adapted Arrhenius model, mass balance, ideal gas law, and Sieverts' law) or numerical fitting (black box model). Such models can be applied for the development of controllers, which is out of the scope of this thesis. A dynamic model of a reforming system (the CMR) is given to simulate the dynamics of the pure hydrogen production rate at unsteady state conditions (between two steady state points) under fuel flow rate and pressure set-up steps.

The model considers the kinetics of the catalytic reforming reactions regarding the molar production of ESR products, especially hydrogen inside the reactor at unsteady operating conditions. Additionally, application of the CMR makes it possible to investigate the effect of the byproducts of the ESR (CO, CO 2 , H 2 O, and CH 4 ) on the performance of a real case Pd-Ag membrane based on the observed reaction kinetics (concentration of the ESR products). The latter is an important factor in monitoring and simulation of the performance of the membrane in ESR environment so that many works have been reported on the investigation of the effect of the gaseous byproducts on the permeation behavior of the membranes.

In this chapter, a static model is developed to calculate the molar production rate of the ESR products (CO, CO 2 , CH 4 , H 2 , and H 2 O) in the CMR. Two different models were developed based on the experimental results to predict the pure hydrogen production rate at steady state conditions. The models are named as model 1 (Sieverts' law) and model 2 (a black box model). Several critical factors such as the partial pressure of hydrogen, the activation energy of the hydrogen permeation process, and most importantly, the effect of co-presence of the ESR products on the permeation activity of the membrane were only taken into account by means of the Sieverts' law, which reinforced and supported the aim of the simulation of the CMR. The aim of developing the black box model (model 2) was to investigate a numerical model not based on the physical rules and definitions (Sieverts' law) that is normally applied for calculation of hydrogen permeation rate. However, the two important factors in hydrogen permeation rate, i.e. fuel flow rate and pressure, were taken into account.

Finally, the alteration of the pure hydrogen production rate (transient state) during the fuel flow rate and pressure changes at isothermal conditions are simulated. The application of the model is in the design and control of a hydrogen producing system capable of online feeding a fuel cell with the possibility of instant adjustments of the pure hydrogen stream, following the changes in the fuel cell electric load.

parameters of the equations. In the formulation of the molar production rate of the ESR products (ṅ ), temperature [K] is noted as 'T', pressure [Pa] as 'P', steam to carbon ratio as 'SC', and the fuel flow rate [m 3 /s] as 'F'.

Accordingly, the fitting parameters were given the values of "0" or "1" to form the various final models. Eight models out of the several tested ones showed interesting results in terms of fitting the experimental (measured via experiments) values. The mentioned models are given in Table 5.1.

Model

Equation on which the model is based parameters ṅ

1 5.1 =1, θ=0, =0 α × F × P × exp( ω × SC + λ RT ) 2 5.1 θ=0, =0 α × F × P β × exp( ω × SC + λ RT ) 3 5.1 =0, θ=1, ω=0 α × F × SC × exp( × P + λ RT ) 4 5.1 =1, θ=1 α × F × P × SC × exp( × P + ω × SC + λ RT ) 5 5.2 =1, θ=0, ω=0 =0 α × F × P × exp( × SC λ RT ) 6 5.2 θ=0, ω=0, =0 α × F × P β × exp( × SC λ RT ) 7 5.2 =1, θ=0, ω=0 α × F × P × exp( × SC λ + RT ) 8 5.2 θ=0, ω=0 α × F × P β × exp( × SC λ + RT )
Table 5.1 General models for calculation of the molar flow rate of the ESR products Apart from the general models, which include all the experimental points (presented in Table 3.1), specific models were developed at specific temperatures, i.e. at 873 K or 923 K, mainly to improve the accuracy of the modeling results. This is a crucial step for the isothermal dynamic simulation of the reformer, which will be discussed in the next section. Indeed, a specific model is a general model, which has been developed only at 873 K or 923 K, hence, one parameter (temperature T) has been omitted in the data fitting, resulting in more precise modeled results (better fitting to the experimental data). The specific models are given in Table 5.2.

Model

ṅ 1 α × F × P × exp( ω × SC RT ) 2 α × F × P β × exp( ω × SC RT ) 3 α × F × SC × exp( × P RT ) 4 α × F × SC × exp( × P ω RT ) 5 α × F β × SC × exp( × P ω + RT )
Table 5.

The specific models

As presented in Table 5.2, the models are functions of fuel flow rate, pressure, and steam to carbon ratio.

The same methodology was used to calculate the permeation rate of hydrogen through the membrane pure hydrogen production rate (pure hydrogen production rate in the permeation zone). Several models were developed as functions of the molar production rate of hydrogen inside the reactor (ṅ ) that was obtained by the steady state model of the catalytic zone, and the operating pressure and temperature. Contributing to the partial pressure of hydrogen inside the reactor, pressure and the rate of hydrogen production are the determining factors in the rate of hydrogen permeation through the membrane. The general form of the hydrogen permeation model is:

= × ṅ × × + 5.4
Where J H2 is the rate of hydrogen permeation [mol/s] through the membrane (permeation zone). 'a', 'b', 'c', 'd', and 'e' are the fitting parameters.

The models based on which the calculated values of the hydrogen permeation rate fitted the experimental ones are given in Table 5.3.

Model

Parameters

1 c=d=0 a × ṅ + e 2 d=0 a × ṅ × P + e 3 b=1, e=0 a × ṅ × P × T 4 e=0 a × ṅ × P × T 5 - a × ṅ × P × T + e Table 5.
3 Models for calculation of the hydrogen permeation rate

Dynamic simulation

The dynamic behavior of the membrane in terms of hydrogen permeation rate at transient conditions was simulated. Transient conditions were applied by acting on the fuel flow rate or operating pressure.

For this purpose, a specific steady state model was used to calculate the total hydrogen production rate in the reactor. In order to monitor the dynamic performance of the reformer, pure hydrogen production (permeation through Pd-Ag membranes) was measured during dynamic tests. Then two different models were developed based on the experimental results to calculate the pure hydrogen production rate at steady state conditions. The models were named as model 1 and model 2, referred to the Sieverts' law and black box models, respectively. The Model 2 presented in Table 5.3 was used as the black box model.

• Dynamic experiments

Using the same experimental setup presented before, several dynamic tests were performed to measure the rate of hydrogen permeation through the membrane at transient conditions.

During the dynamic experiments, the fuel flow rate and the operating pressure of the reactor were altered. While giving time to the system to reach the next steady state point in terms of pure hydrogen flow rate, the variation of the hydrogen permeation rate over time was recorded. Therefore, the reformer can be assumed to have two different controlling parts at isothermal conditions: the fuel flow rate controller, and the pressure controller. A schematic plan of the fuel reformer system is shown in Fig. 5.2 to illustrate the fuel flow rate and pressure control subdivisions of the reformer. It is assumed that the fuel (ethanol+water) is in its gas phase at the entrance of the volume of the CMR and the ideal gas law is applied.

CMR isothermal simulation

• Steady state model of the catalytic zone

The Model 5 presented in Table 5.2 for the catalytic zone was used to calculate the total molar production rate of species present in the catalytic zone of the CMR, i.e. CO, CO 2 , CH 4 , H 2 , and H 2 O as the products of the catalytic conversion of ethanol (around the membrane). At constant steam to carbon ratio and ω=1:

ṅ = × × exp( -( × + ) ) 5.5
F [m 3 /s] and P [Pa] represent the fuel (ethanol + water) flow rate and the reactor pressure, respectively. The molar production rate of the ESR products were used to calculate the partial pressure of hydrogen (P H2,r in eq. 5.6) around the membare surface (right before the permeation zone).

• Steady state models of the permeation zone

The hydrogen permeation rate through the membrane at steady state was modeled via two different pathways. The first model, named as "model 1", was derived from the Sieverts' law via which the hydrogen permeation phenomenon through the membrane is explained based on the mass transfer and surface reactions principals. As stated by the Sieverts' law, hydrogen permeation is a temperature activated phenomena driven by the difference between the partial pressure of hydrogen at two sides, i.e. the retentate side (inside the reactor, around the membrane) and the permeate side (right after the membrane) [START_REF] López | Hydrogen production from ethanol over Pd-Rh/CeO2 with a metallic membrane reactor[END_REF][START_REF] Domínguez | Ethanol steam reforming at very low temperature over cobalt talc in a membrane reactor[END_REF]:

= δ - ( � , -� , ) 5.6
Where J is the pure hydrogen production rate obtained via the Sieverts' law (model 1). 'δ' is the thickness of the membrane and Q 0 is the pre-exponential factor. E a , R, and T are the activation energy, the universal gas constant, and temperature, respectively. P H2,r and P H2,perm are the partial pressure of hydrogen at the retentate and permeate sides, respectively. The partial pressure of hydrogen inside the reactor was calculated based on the hydrogen fraction in the gas phase assuming that the only present species in the catalytic bed (and around the membrane) are CO, CO 2 , CH 4 , H 2 and H 2 O. Therefore: , = × , 5.7

Where P and y H2,r represent the reactor pressure and the fraction of hydrogen in the catalyst bed, respectively. The activation energy (E a ) and the pre-exponential factor (Q 0 ) are calculated by means of permeation experiments during which pure hydrogen at known temperature and pressure is purged and the permeation rate of hydrogen through the membrane is measured (atmospheric pressure at the permeate side) [START_REF] Hla | Modelling and experimental studies of a watergas shift catalytic membrane reactor[END_REF][START_REF] Basile | Water gas shift reaction in membrane reactors: Theoretical investigation by artificial neural networks model and experimental validation[END_REF][START_REF] Sanz | Hydrogen production in a Pore-Plated Pd-membrane reactor: Experimental analysis and model validation for the Water Gas Shift reaction[END_REF][START_REF] Chein | Three-dimensional numerical modeling on high pressure membrane reactors for high temperature water-gas shift reaction[END_REF][START_REF] Marcoberardino | Fixed bed membrane reactor for hydrogen production from steam methane reforming: Experimental and modeling approach[END_REF][START_REF] Iulianelli | Model biogas steam reforming in a thin Pd-supported membrane reactor to generate clean hydrogen for fuel cells[END_REF].

As discussed in chapter 2, the published open literature offers no robust model or pattern on the effect of different species on the performance of the membrane in the real atmosphere of methane steam reforming and water gas shift reactions. It was concluded that to understand the influence of co-existence of the ESR products on the permeation performance of the membrane, special models must be developed regarding specific operational conditions of the ESR environment.

Accordingly, a model was developed for hydrogen permeation through the Pd-Ag membrane; specifically for the ESR environment at the operating conditions presented in this work. It is assumed that the concentrations of CO and H 2 O affect the permeation P, V, T, and M W are reactor pressure, the volume of the reactor, reactor temperature, and the molar mass of the fuel mixture, respectively. 'm' is the accumulated mass of the fuel added to the reactor volume. It was assumed that the accumulation rate of the pumped fuel into the constant volume of the reactor at constant temperature, results in pressure rise as the pressure valve acts on the outlet of the system to block the retentate stream when pressure increase is required. Conversely, the pressure valve lets the retentate gas be released to decrease the reactor pressure, so that the inlet mass flow rate of the fuel gets lower than the outlet mass flow rate. Regardless of the action of the pressure valve on the retentate gas stream, hydrogen constantly permeates through the membrane. Therefore, the added mass to the reactor volume is the difference between the fuel flow rate, and the retentate gas flow rate plus hydrogen permeation rate, so that:

= ṁ -ṁ -ṁ .

5.11

Where ṁ fuel and ṁ H2.perm represent the fuel flow rate and hydrogen permeation rate, respectively, both in [kg/s]. Then, the ideal gas law is written as:

= ( ) ×

5.12

Where is the rate of the accumulation of the mass in the reactor volume. In this work, the CMR is a packed bed reactor running at isothermal conditions, with negligible axial mixing. The temperature and concentration difference is neglected, so that the models are considered as ideal plug flow pseudo-homogenous ones [START_REF] Jakobsen | Chemical Reactor Modeling: Multiphase Reactive Flows[END_REF].

To develop the dynamic model of the permeate zone, a first order function was used:

= , + 5.13
is the pure hydrogen production rate obtained by the dynamic model. The superscript "D" stands for the dynamic model. J , represents the hydrogen permeation rate calculated via model 1 or 2, considering every single operating point as steady state. The time constant is presented as τ. The measured dynamic of fuel flow rate was faster than the sampling time (<1 second). Therefore, the fuel flow rate is always equal to its set point value:

= 5.14
Finally, equation 5.13 is written as:

= , + 5.15
Where F is the fuel flow rate set point (see Fig. 5.4).

The simulation was performed by means of Ordinary Differential Equation (O.D.E) solver.

The dynamic simulation of the reforming system was performed in Matlab Simulink. The Simulink simulator used to model the reactor pressure and pure hydrogen flow rate is shown in Fig. 5.5. The simulated pressure is used as an input for the calculation of the partial pressure of hydrogen and the production rate of pure hydrogen via the Sieverts' law. The set point pressure value is needed since a PID controller is used to set the pressure and act on the pressure valve.

Results and discussion

Least Square Method (LSM) was applied to obtain all the fitting parameters regarding the steady state models. The time constant was estimated from a set of trials and errors. The time constant is a function of operating conditions (pressure or fuel flow rate). This is attributed to the dependency of the time constant on the function of the CMR during pressure or fuel flow rate change cycles. The response of the system to the fuel flow rate adjustments was faster due to instant modification made by the pump and instant increase in the inlet flow rate of the fuel into the reactor. In the case of pressure change steps, the CMR needed to be given time to increase the reactor pressure gradually as the fuel is added to the reactor constant volume.

Steady state models of the permeation zone

The products of ESR (CO, CO 2 , CH 4 , H 2 , and H 2 O) and the pure hydrogen permeation rate was modeled at four different fuel flow rates, i.e. 50, 100, 150, 200 µl/min and three different pressures (6, 8, and 10 bar). As mentioned before, the molar production rate of all the ESR products is needed in order to calculate the partial pressure of hydrogen in the

The modeled values could fit the experimental results within the 10% error, especially in the case of production rate of hydrogen. The values of the fitting parameters (eq. 5.5) for all the gases are given in Table 5 According to eq. 5.5, it can be seen that the values of ×P are very small compared to , except in the case of hydrogen. The most effective factor on the hydrogen permeation is the partial pressure of hydrogen in the reactor; hence, the value of ×P is higher in this case. The same explanation can be given regarding the parameter . In the case of hydrogen, the effect of pressure in the CMR configuration is dominant in comparison with the fuel flow rate, resulting in the smallest value of in the case of hydrogen. Conversely, the value of in the case of CH 4 is the highest among the gases because the only source of CH 4 is the ethanol decomposition reaction (eq. 2.1). At complete ethanol conversion, the higher the fuel flow rate is, the higher the production rate of CH 4 is. The value of in the case of H 2 O is nearly one, which is very relevant since the ESR reaction were performed at SC=3 where there is a large amount of excess water. It can be concluded that the molar flow rate of water in the retentate stream is proportional to the inlet molar flow rate of water in the fuel mixture (ethanol + water). At SC=3, a large portion of the inlet water (70-90%) leaves the reactor in the form of steam as unreacted water. The value of in the case of CO is one order of magnitude smaller than other gases, which is attributed to the very small amount of CO detected at the outlet of the reactor. The values of proves that at higher pressures, less byproducts (CO, CO 2 , CH 4 , and H 2 O) and more hydrogen are generated, which is totally in agreement with the experimental results and the aim of the application of the CMR, where ESR reaction is promoted (the shift effect).

The results of the modeling of the molar flow rate of the ESR products by all the formulations presented in Table 5.1 and Table 5.2 (general models and specific models, respectively) are given in appendix B.

The result of the pre-exponential factor model (eq. 5.9) shows that the model fits very well to the calculated values except at P=6 bar, so that the coefficient of determination (R The fitting parameters of the model 2 (a, b, c, and e) do not possess a physical meaning, but the value of the parameter 'c' proves the noticeable effect of pressure on the hydrogen permeation rate. Besides, as the value of the parameter 'e' is very small, it is shown that at the pressure or fuel flow rate of zero, almost no hydrogen is generated.

Isothermal Dynamic simulation

• Pressure change simulation Keeping in mind the configuration of the CMR, when the pressure of the reactor is set at a higher value, the outlet of the reactor is blocked, the flow rate of the retentate gas (ṁ ) is zero (see eq. 5.10), so that the inlet fuel is added to the volume of the reactor to increase the pressure. On the contrary, when reactor pressure is set at a lower value, the pressure valve is opened so that gas is released and reactor pressure drops rapidly. The different behavior of the system during pressure increasing and decreasing steps is due to the different act of the pressure controlling system on the pressure valve (see Fig. 5.2). Therefore, the dynamics of the pressure control differs in different steps. The importance of such a performance lies in the dependency of pure hydrogen permeation rate through the membrane on the partial pressure of hydrogen in the reactor. The simulated pressure change behavior of the reformer system is shown in Fig. 5.10.

conditions. The black box model is a simulation based on pure fittings, which may be totally dependent on certain operating conditions.

Conclusion

Ethanol steam reforming (ESR) over Pd-Rh/CeO 2 catalyst was performed in a CMR unit at 923 K, 6-10 bar, and fuel flow rates of 50 to 200 µl/min using a mixture of ethanol and distilled water with steam to carbon ratio of Finally, the dynamic simulation was performed under ideal gas law assumptions to simulate the permeation behavior of the membrane and the dynamics of the pure hydrogen production rate (permeation zone) in the case of the fuel flow rate or the operational pressure set point adjustment at isothermal conditions. The simulation results based on both model 1 and model 2 fitted the experimental observations. The effective critical factors such as hydrogen partial pressure in the CMR and the influence of the co-existence of the ESR products on the permeation behavior of the membrane were taken into account by the Sieverts' law model, presenting a robust simulation. Although the pressure change steps simulation results were more satisfactory, the simulation of the dynamics of the fuel flow rate change is more essential, as the system responds much faster to such an adjustment. The presented dynamic models resembled hydrogen flow rate adjustments needed to follow changes in the electric load of a fuel cell when fed online, which is an essential step to design, control, and optimize an ethanol to electricity system.

156

In this work, a comprehensive investigation of production of fuel cell grade hydrogen (pure hydrogen) in a catalytic membrane reactor (CMR) via ethanol steam reforming (ESR) over a Rh 0.5 -Pd 0.5 /CeO 2 catalyst was given.

The highlights of the experimental investigation of the ESR in the CMR are summarized as:

• The ESR experiments at several operating conditions showed high selectivity of the catalyst towards H 2 (>50% volume) together with very low molar production rate of CO per mol of ethanol in the feed. Besides, complete ethanol conversion was reached at all the operating conditions.

• The shift effect resulted in significantly higher hydrogen production, thanks to the especial configuration of the CMR where the membrane was fully covered by the catalyst pellets. The equilibrium limitations were conquered at high pressure as a result of the shift effect, leading to a positive effect of pressure on the performance of the CMR despite unfavorable thermodynamic conditions at high pressure.

• The influence of species present in the CMR such as CO and CH 4 was shown based on the experimental results of the permeation tests in the atmosphere of pure hydrogen and ethanol steam reforming.

• A comparison between the CMR and the SMR (staged membrane reactor) clearly revealed the superior performance of the CMR. Thanks to the special configuration of the CMR and brilliant reactivity of the catalyst, pure hydrogen production rate and methane conversion were promoted at high pressure in the case of the CMR.

• 0.9 L N pure hydrogen/ml EtOH at 12 bar and 923 K was obtained, equal to 3.2 mol of pure hydrogen per mol of inlet EtOH. A hydrogen yield of 0.55 was reached at 12 bar and 923 K, much higher than that of SMR, 0.2. Hydrogen recovery reached values up to 95% in the CMR, significantly higher than 70% in the case of the SMR.

The results of the thermodynamic analysis are as follow:

• Exergy analysis was introduced as a relevant tool to explain the effects of operating conditions and thermodynamic losses on both pure hydrogen production rate and total efficiency of the membrane reactor.

• A significant effect of pressure and temperature on the exergy efficiency was observed.

Following the pure hydrogen production rate, the exergy efficiency increased with pressure because of the shift effect, where more methane was converted at higher pressure resulting in higher production rate of pure hydrogen.

• Exergy analysis showed that the highest exergy destruction occurred via the retentate gas stream and heat losses. The exergy content of the retentate gas especially at lower pressures could provide the reactor with a notable fraction of its required heat at steady state conditions. This would result in a notable increase of the exergy efficiency of the system.

• By using the retentate gas in an insulated reactor, the exergy efficiency was increased drastically and placed between 70-90%. This result was expected since, in one hand, exergy destruction decreased and, on the other hand, the energy requirement of the system was partially or totally met by utilization of the retentate gas. In the case of an insulated reactor, exergy efficiency up to around 50% was reached at 12 bar and 923 K.

In the case of recovery of the retentate gas and insulation of the reactor, exergy efficiency placed between 70-90%.

• Based on the exergy evaluation, the best operating conditions were calculated at the highest pressure and temperature used in this work (923 K and 12 bar) and the lowest S/C ratio of 1.5, where carbon deposition is probable.

• Thermal efficiency of the CMR was between 60-90% for an insulated reactor system and decreased to around 40-60% when the heat loss was considered. The consideration of the thermal efficiency revealed that while thermal efficiency offers an ideal performance at high pressure and temperature, exergy efficiency discloses inevitable irreversibilities even in the case of an insulated reactor.

• In the case of the SMR, at least 80% of the inlet exergy was lost by heat losses and the retentate stream. The insulation of the SMR resulted in a significant increase of the thermal efficiency, but not exergy efficiency since a large amount of methane and unpermeated hydrogen left the reactor as unused exergy.

• In the case of the recovery of the retentate gas in an insulated reactor, the exergy efficiency of the SMR was slightly higher than the CMR, however, much higher pure hydrogen yield and production rate is obtained in the case of the CMR.

Finally, the outcomes of the dynamic simulation of the CMR are:

• A static model for the catalytic zone was derived from the Arrhenius law as a function of operating conditions based on the experimental results to simulate the total molar production rates of ESR products. The results of the model fitted the experimental value very well (coefficient of determination R 2 = 0.993).

• The permeation zone (membrane) was modeled based on the Sieverts' law as the physical definition of hydrogen permeation through the membrane via which the robustness of the simulation was proved (R 2 = 0.86). Besides, a black box model as a function of the fuel flow rate and the operating pressure was developed and successfully fitted to the experimental values (R 2 = 0.91).

• A dynamic model was proposed under ideal gas law assumptions to simulate the dynamics of pure hydrogen production rate in the case of the fuel flow rate or the operational pressure set point adjustments (transient state) at isothermal conditions. Both pressure and fuel flow rate change steps simulations fitted the experimental values very well, proving a successful essential step needed to simulate an "Ethanol to Electricity" system. The simulation of the dynamics of the fuel flow rate change was more essential, as the system responds much faster to such an adjustment.

• The results of the dynamic simulation based on the Sieverts' law proved the robustness of the simulation and reinforced and supported the aim of the simulation of the CMR.

Several critical factors such as the partial pressure of hydrogen, the activation energy of the hydrogen permeation process, and most importantly, the effect of co-presence of the ESR products on the permeation activity of the membrane were taken into account by means of the Sieverts' law.

• Future avenues

A comprehensive experimental and thermodynamic study of pure hydrogen producing system was given in this PhD thesis. However, there are promising opportunities to continue the presented work and fulfil the application of the studied system.

Future works can be devoted to the study of different noble based catalysts in the CMR. The activity of different catalysts in terms of ethanol conversion, hydrogen selectivity, and hydrogen yield and recovery can be studied similar to what reported in this thesis, to compare different catalysts in the studied membrane reactor in wider ranges of the operating conditions.

Besides, the exergy evaluation data presented in this thesis can be used for the exergonomic evaluation of the reforming system, which is an essential and modern way to scale up. Additionally, life cycle assessment of such a reforming system, especially in connection to a fuel cell is a promising method to give a comprehensive idea on the environmental impact of the studied reforming system in this thesis. The mentioned methods (exergonomic and life cycle assessment) may add additional value to the introduction and commercialization of the reforming system.

In the case of the retentate flow rate measurement, for every experimental point (certain temperature, pressure, S/C ratio, and fuel flow rate), 5 measurements were carried out and the arithmetic mean value of the measurements was used for experimental evaluation purposes:

� = ∑ =

A2

X � is the arithmetic mean of the 5 measurements (n=5) of X j (retentate flow rate). Then, the variance of the probability distribution of X j is:

( ) = - ∑ ( = -� ) 2 A3
The variance of the mean is given as:

( � ) = ( )

A4

Finally, the standard uncertainty is obtained:

( ) = ( � ) A5
If an evaluation equation (y) is considered as a function of the independent measured factors:

Y = f(x 1 , x 2 , x 3 ,…x n ) A6
The standard combined uncertainty (u c (y)) is defined as:

() = ∑ ( = ) ( ) A7
This method is used to calculate the uncertainty of several experimental and thermodynamic factors that are calculated by different formulations. The detailed analysis of the uncertainty analysis of the experimental and thermodynamic factors is given in the following sections.

A.1 The uncertainty of the experimental evaluation factors

Y = H . × OH Hydrogen yield A8 R = H .

H . a

Hydrogen recovery A9

Ө = H .
H .

Pure hydrogen efficiency A10

'F' is the molar flow rate in mol/s. F H2.total represents the total hydrogen production rate via ESR in the CMR. Inlet hydrogen consists of the hydrogen content of the fuel mixture (water + ethanol). Therefore, the uncertainty of each equation is given:

A.1.1 Hydrogen yield (Y ) = � H . � �F . � + � OH � (F O ) A11 H . = × OH A12 OH = -H . ×( OH )

A13

The uncertainty of F . is the same as the uncertainty of the pure hydrogen mass flow measurement. ) (F . ) A18

A.1.2 Hydrogen recovery

R =

F . is the flow rate of unpermeated hydrogen which appears in the retentate stream. u(F H . ) Itself is obtained separately through a similar method. The molar flow rate of CO 2 , H 2 , CH 4 , and CO in the retentate stream is a function of the selectivity analysis done by the GC and the flow rate of the retentate gas measured by human by using a bubble meter. If the flow rate of the retentate is in ml/min and the selectivity of the retentate gas in percentage, then the molar flow rate of any gaseous product in mol/s is calculated as:

F = × × A19
A is a constant value for each gas for the unit conversion, to obtain the flow rate of each gas in mol/s. F represents the retentate flow rate (ml/min). S GC is the selectivity of gas, in percent (0<S GC <100). Knowing the uncertainty of the F and S C :

�F � = ( a ) ( ) + ( a ) ( ) A20 �F � = ( ) + ( ) A21
A.1.3 Pure hydrogen efficiency

�Ө � = � H . � �F . � + � H � (F ) A22 H . = H A23 H = -H . ( H ) A24 F = F O + 3 × F O A25 (F ) = (F O ) + 9 × (F O ) A26
'F' is the molar flow rate in mol/s.

A.1.4 The uncertainty of the fuel blend (S/C ratio) and molar fuel flow rate

The fuel is a mixture of distilled water and pure ethanol. It is assumed that water and ethanol are pure. A graduated cylinder is used for measurement of the required volume of each component for fuel preparation. The accuracy of the graduated cylinder is ±0.1 ml. A constant human error of ±0.5 ml is considered. Therefore, the total maximum error of ±0.6 ml is considered in the measurement of the required volume of water and ethanol. This is equal to 0.033 and 0.010 mole of water (n O ) and ethanol (n O ), respectively. The uncertainty of the S/C ratio is calculated as:

/ = × A27
n represents the number of the moles:

(S/C) = ( ( /C) ) ( ) + ( ( /C) ) ( ) A28 ( /C) = × A29 ( /C) = - × A30
The experiments were performed at two fuel flow rates i.e. FF=50 l/min (0.05 ml/min) and FF=100 l/min (0.1 ml/min). The uncertainty of the molar flow rate of water and ethanol because of the errors in fuel preparation is calculated as:

u(n O ) = . × A31 u(n O ) = . ×

A32

V mixture is the volume of the mixture of the ethanol and water. At each S/C ratio, 100 ml ethanol is mixed with the relevant volume of water, which is 187, 123.6, and 100 ml of water to make a mixture with S/C ratio of 3, 2, and 1.6, respectively. Therefore, V mixture is equal to 287, 223.6, and 200 ml at S/C ratio of 3, 2, and 1.6, respectively.

The pump used in the experimental setup showed an error equal to 5% in the flow rate of the fuel, which means 2.5 and 5 l/min at FF=50 and 100 l/min, respectively. It is assumed that the error of the pump stays unchanged at higher pressures (maximum pressure of 12 bar was reached in this work).

Finally, the maximum possible uncertainty of the molar flow rate of ethanol and water was considered to be the sum of uncertainty of fuel blend preparation and the uncertainty of the pump (u(n O ) and u(n O ) ):

u(n O ) = u(n O ) + u(n O ) A33 u(n O ) = u(n O ) + u(n O ) A34 A.1.

Steam in the reactor (unreacted water) S reactor

Unreacted steam in the reactor was used in the estimation of the hydrogen fraction in the gases inside the reactor (around the membrane) to be used in the Sieverts' law equation. The unreacted steam in the reactor (S reactor ) is supposed to be difference between the inlet steam (S inlet ) and the amount of steam used in the reforming reactions (S used ):

= -A35

The uncertainty of the inlet flow rate of steam is the same as inlet water in the fuel:

( ) = u(̇ O ) A36
Steam is used in water gas shift reaction WGSR (S ) and methane steam reforming (S ) reactions:

= + A37
According to the ESR reactions mechanism (eq. 2.2 and 2.3), total rate of steam consumption is equal to the rate of CO consumption plus twice of the rate of CH 4 consumption. For each mole of decomposed ethanol, one mole of methane and one mole of carbon monoxide are formed. Therefore:

= = ̇ - A38 = 2 × = 2 × (̇ -̇ ) A39
Accordingly:

= 3 × ̇ --2 × A40 ( ) = 3 × (̇ ) + ( ) + 2 × ( ) A41
Finally:

( ) = ( ) + ( ) A42

A.2 Uncertainty analysis of the exergy evaluation

To introduce the formulation of uncertainty analysis of the exergetic evaluation, some exergy terms are recalled:

= - A43 = - A44 = 1 - A45 A.2

.1 Inlet exergy

= + + A46

W el represents the electrical exergy rate, which is assumed equal to the exergy rate needed for the ESR plus the heat losses.

A.2.2 Uncertainty of the exergy rates of water ( ) and ethanol ( )

The exergy of every species consists of physical ( ℎ ), chemical ( ℎ ), and mixing exergy ( ):

= ℎ + ℎ + A47
Mixing exergy is very small in front of the chemical and physical exergy ( a < 0.005), therefore, this term is neglected in uncertainty analysis calculations:

( ) = � ℎ � + ( ℎ ) A48
Uncertainty of chemical exergy is easily calculated since the thermodynamic tables were used to obtain the standard chemical exergy of the species:

EX = ℎ × ̇ A49 u(EX ) = ex × u(ṅ ) A50
Physical exergy is obtained as a function of the specific enthalpy and entropy:

ℎ = h -h -T (s -s ) A51
h and s are the specific enthalpy and entropy at reaction conditions (T r ). h 0 and s 0 (specific enthalpy and entropy at reaction conditions) are considered fixed without error since the reference temperature is assumed to be 298 K. NASA polynomials are used to calculate the specific enthalpy and entropy:

= a + a + a + a + a + 6 A52 = a Ln(T) + a T + a + a + a + a A53
Then:

∂ ∂ = R × (a + a T + a T + a T + a T ) A54 ∂ ∂ = R × ( + a + a T + a T + a T ) A55 u(h) = ∂ ∂ × u(T ) A56 u(S) = ∂ ∂ × u(T ) A57 u �ex � = u (h) + T × u (s) A58 EX = ex × ṅ A59 u �EX � = ( ∂ a ∂ a a ) × u �ex � + ( ∂ a ∂ ̇ ) × u (ṅ ) A60 = ṅ × u �ex � + ex × u (ṅ ) A61
This formulation is also true for the physical and chemical exergy content of the components of the retentate gas. For ethanol and water, the physical exergy is assumed zero, since they enter the system at reference conditions.

A.2.3

Uncertainty of electrical energy consumption (in the form of exergy) consists of three terms, i.e. the required exergy for fuel evaporation and heat up to required temperature (T r ), required heat for the ESR reactions, and heat losses.

= + + A62 � � = � � + � � + � � A63 A.2.

Uncertainty of

The reaction promotion of WGS and MSR are measured by means of the consumption rate (reaction) of CO and CH 4 , respectively. Therefore, the required heat for ESR is obtained based on the promotion of WGS and MSR reactions.

= × ̇ + × ̇ A64
ΔH represents the enthalpy change of the reaction.

� � = × �̇ � + × �̇ � A65 �̇ � = (̇ ) + (̇ ) A66 �̇ � = (̇ ) + ( ) A67
A.2.5 Uncertainty of Q stands for the required exergy needed for the evaporation of the water (ΔH O ) and ethanol (ΔH O ) and to heat the vapors up to the reaction temperature (T r ).

= ΔH O × ṅ O + ΔH O × ṅ O A68 u � � = ΔH O × u(ṅ O ) + ΔH O × u(ṅ O ) A69

A.2.6 Uncertainty of

The reactor vessel is a stainless still cylinder via which the heat is transferred (lost) to the reference environment.

EX = π ( - a ) ( ) A70
'L' is the length of the tube (reactor) and 'K' stands for the thermal conductivity of the insulation, which is glass wool. T wall and T r represent the reactor vessel wall temperature and the reaction (reactor) temperature, respectively.

�EX

� = π ( )

× ((T ) + (T )) A71

A.2.7 Uncertainty in temperature reading:

The reactor (reaction) temperature is measured by means of a K-type thermocouple. The same method was used to measure the wall temperature. The temperature measurement equipment has the following errors:

• Thermometer: 0.3% reading+1℃ digit • K-type thermocouple: 0.75% of the read value (Reference: the data base/data sheet of the product)

Based on the mentioned error, uncertainty of T wall and T r and two different temperatures operated in this work (873 and 923 K) is calculated:

• (T ) T (at T = 873 K) =

A.2.8 Outlet exergy

The outlet exergy rate consists of the exergy rate of the retentate gas and the exergy rate of the pure hydrogen stream permeating through the membrane:

A.3 The results of the uncertainty analysis of the experimental evaluation factors

The uncertainty of the pure hydrogen mass flow meter is 5.77×10 -7 mol/s or 0.87 ml/min, regardless of the operating conditions. The value of is positive only in the case of hydrogen, which proves that the production rate of hydrogen in the CMR is directly proportional to the operating pressure. The values of ω in the case of CO 2 and H 2 O are positive, proving that at higher S/C ratio, more CO 2 is formed, and naturally, there is excess water fed into and leaving the reactor. The smallest value of ω belongs to CO that has the lowest selectivity among the ESR products. Very high values of can be ascribed to the significant effect of the temperature on the ESR reaction promotion and hydrogen permeation rate.

In the case of H 2 , the value of is very small, but on the other hand the value of α is much higher compared to other ESR product. Considering the formulation of this model, the difference between the values of and α is interpreted as the significant effect of pressure on the hydrogen production rate and the minor effect of S/C ratio. Besides, the very large value of in the case of hydrogen, represents the effect of temperature.

The value of is larger in the case of hydrogen, which proves that the production rate of hydrogen in the CMR is directly proportional to the operating pressure. Very large values of specially compared to can be explained by the significant effect of temperature in comparison with the S/C ratio. The value of is near 1 in the case of H 2 O since the amount of eater is proportional to the S/C ratio. Very high values of can be ascribed to the significant effect of the temperature on the ESR reaction promotion and hydrogen permeation rate.

H 2 O has the highest value of ω that is attributed to the effect of S/C ratio on the production rate of H 2 O. The amount of excess water increased with S/C ratio. Comparatively, the value of ω in the case of hydrogen presents the minor effect of S/C ratio on the production rate of H 2 . The positive values of ω in the case of CO 2 and H 2 O can be attributed to the fact that the ESR reaction is promoted better at higher S/C ratio, probably due to the availability of excess water. The highest value of α belongs to H 2 , proving the effect of pressure on the hydrogen production rate.

The value of is larger in the case of hydrogen, which proves that the production rate of hydrogen in the CMR is directly proportional to the operating pressure. The positive values of ω in the case of CO 2 and H 2 O can be attributed to the fact that the ESR reaction is promoted better at higher S/C ratio, probably due to the availability of excess water.

The positive and larger value of in the case of hydrogen proves that the production rate of hydrogen in the CMR is directly proportional to the operating pressure. The positive values of θ in the case of CO 2 and H 2 O can be attributed to the fact that the ESR reaction is promoted better at higher S/C ratio, probably due to the availability of excess water.

Résumé

1 3

 13 Universitat Politècnica de Catalunya and The Department of Energy Systems and Environment in Ecole des Mines de Nantes for creating a pleasant working environment. Catalytic conversion of ethanol 2.2.2 Side products of reforming processes 2.3 Ethanol steam reforming in the literature 2.3.1 Ethanol Steam Reforming over catalysts without noble metals 2.3.2 Ethanol Steam Reforming over noble metal-based catalysts 2.4 Catalytic Membrane Reactors (CMRs) 2.4.1 Ethanol steam reforming in CMRs 2.4.2 Effect of co-existence of ESR products on the permeation behavior of Pd-Ag membrane Chapter 3: Experimental results of ethanol steam reforming over Pd-Rh/CeO 2 in the Pd-Ag membrane reactor 3Comparison between the CMR and the SMR configurations 3.3.1 Overall performance of the CMR and SMR 3.3.2 The effect of pressure 3.3.3 Pure hydrogen efficiency 3.4 Conclusion Chapter 4: Thermodynamic analysis of the CMR and SMR systems 4.1 Exergy analysis 4.2 Materials and methods 4.3 Results and Discussion 4.3.1 Thermodynamic analysis of the CMR • Effect of operational conditions • Efficiency improvement • Retentate gas recovery • Insulation of the CMR 4.3.2 Thermodynamic analysis of the Stage Membrane Reactor (SMR) • Efficiency improvement 4.3.3 Comparison between the CMR and SMR configurations 4.4 Conclusion Chapter 5: Modeling and dynamic simulation of the reforming system 5.1 Dynamic simulation of the CMRs 5.2 Materials and methods 5.2.1 General and specific models 5.2.2 Dynamic simulation • Dynamic experiments 5.2.3 CMR isothermal simulation • Steady state model of the catalytic zone • Steady state models of the permeation zone • Isothermal dynamic simulation of the permeation zone 5.3 Results and discussion 5.3.1 Steady state models of the permeation zone 5.3.2 Isothermal Dynamic simulation • Pressure change simulation • Fuel flow rate change simulation 5

Fig. 19 :

 19 Fig. 19: Schéma simulink A T=293 K, des échelons de pression ont été effectués (SC=3, FF=200 l/min), voir Fig. 20.

Fig. 20 :

 20 Fig. 20: Comparaison mesure-modèle lors d'un échelon de pression : pression Lors d'un échelon positif, à débit constant, la pression change car la vanne pilotant le débit des gaz rétentat est fermée. Le bilan masse est donc = ṁ -ṁ . . La pression croît de manière linéaire. Par contre, lors d'un échelon négatif, la vanne pilotant le rétentat s'ouvre complétement et la pression varie très rapidement.

Fig. 2 . 2

 22 Fig. 2.2 Scheme of a catalytic membrane reactor Fig. 2.3 Effect of gas mixtures on the permeation performance of the Pd membrane at 623 K. Figure taken from the work by Sanz et al. [168].

Fig. 3 . 1

 31 Fig. 3.1 Ceramic monolith [41] Fig. 3.2 Prepared catalyst (left hand side) and its microscopic picture (SEM) (right hand side) Fig. 3.3 Scheme of the fuel reformer Fig. 3.4 Scheme of the reactor with the catalyst Fig. 3.5 CMR module Fig. 3.6 Experimental setup Fig. 3.7 The calibration curve of the pump Fig. 3.8 Dependence of hydrogen transport through the Pd-alloy membrane on pressure and temperature. Figure taken from the work by Papadias et al. [143].

Fig. 3 . 9 Fig. 3 .Fig. 3 . 6 Fig. 3 .

 393363 Fig. 3.9 Scheme of the configurations of the reactors: Catalytic Membrane Reactor (CMR), and Stage Membrane Reactor (SMR) Fig. 3.10 Selectivity of the ESR products at atmospheric pressure Fig. 3.11 Moles of methane produced in the CMR per mol of ethanol in the feed Fig. 3.12 Total and permeated moles of hydrogen per mol of ethanol in the feed Fig. 3.13 Molar production rate of CO per mol of ethanol in the feed Fig. 3.14 Hydrogen permeation rate at different operating conditions Fig. 3.15 Hydrogen yield Fig. 3.16 Hydrogen recovery Fig. 3.17 Hydrogen content in the retentate side and total hydrogen production as a function of pressure at 923 K

Fig. 3 .Fig. 4 . 4 Fig. 4 .Fig. 4 .Fig. 4 .Fig. 4 .Fig. 4 . 6 Fig. 4 . 6 Fig. 4 .

 344444446464 Fig. 3.24 Molar flow rate of CH 4 and CO at the reactor exit for CMR and SR configurations Fig. 3.25 Retentate flow rate Fig. 3.26 Pure hydrogen production rate per mole of ethanol in the feed Fig. 3.27 Pure hydrogen production efficiency at stoichiometric conditions Fig. 4.1 Cross sectional scheme of the reactor vessel Fig. 4.2 Scheme of the reformer system Fig. 4.3 Thermal efficiency (a) and exergy efficiency (b) versus pure hydrogen permeation rate Fig. 4.4 Effect of pressure on exergy efficiency at 873 and 923 K Fig. 4.5 Effect of S/C ratio on exergy efficiency Fig. 4.6 Thermal efficiency of the CMR Fig. 4.7 Thermal efficiency (only pure hydrogen) versus pure hydrogen permeation rate Fig. 4.8 Thermal efficiency of the CMR when only pure hydrogen is considered Fig. 4.9 Exergy flows and the share of the different components in the inlet and outlet streams at 923 K and 12 bar Fig. 4.10 EX retentate /EX in and EX H2.pure /EX in vs the pure hydrogen production rate Fig. 4.11 Exergy rate of the retentate gas and required exergy rate over the exergy rate of the inlet vs. pressure at T = 923 K Fig. 4.12 Exergy rate of the retentate gas and required exergy rate over the exergy rate of the inlet vs. pressure at T = 873 K

Fig. 5 . 3

 53 Fig. 5.3 Scheme of the pressure change for the dynamic tests.

Fig. 5 . 4 Fig. 5 . 6

 5456 Fig. 5.4 Scheme of the fuel flow rate change for the dynamic tests. a) Steps of 50 l/min, b) steps of β5 l/min Fig. 5.5 Simulink model to simulate the reactor pressure and the pure hydrogen production rate Fig. 5.6 Parity plots of the ESR products calculated by the static model (eq. 5.5).

Fig. 5 . 7

 57 Fig. 5.7 Result of the pre-exponential factor model (eq. 5.9) when P=6 bar is included (a) and is not included (b) Fig. 5.8 Parity plots of the hydrogen permeation rate obtained by the Sieverts' law model (model 1) when P=6 bar is included (a) and is not included (b) Fig. 5.9 Parity plots of the hydrogen permeation rate obtained by the black box model (model 2) Fig. 5.10 Measured and simulated reactor pressures in the pressure change dynamic tests. T=λβγ K, FF=β00 l/min.

Fig. 5 .

 5 Fig. 5.11 Sieverts' law simulation (model 1) results of pure hydrogen production for pressure change dynamic tests. T=λβγ K, FF=β00 l/min.

Fig. 5 .

 5 Fig. 5.12 Black box (model 2) simulation results for pressure change dynamic tests. T=λβγ K, FF=β00 l/min.

Fig. 5 .

 5 Fig. 5.13 Sieverts' law simulation (model 1) results of pure hydrogen production for fuel flow rate change dynamic tests. a) Steps of 50 l/min, b) steps of β5 l/min (P=10 bar) Fig. 5.14 Black box (model 2) simulation pure hydrogen production for fuel flow rate change dynamic tests. a) Steps of 50 l/min, b) steps of β5 l/min (P=10 bar)

Fig. 1 .

 1 Fig. 1.1 CO 2 emission by sector in EU [1] As determined by EU, to meet the electricity requirements together with clean energy production targets, the share of renewable resources in power generation should be 35, 43, and 50% in 2020, 2030, and 2050, respectively. Particularly in the electricity generation sector, the share of renewable electricity must reach 45% in 2030 and 53% in 2050. The net electricity production by the type of fuel for EU is given in Fig. 1.2.

Fig. 1 . 2

 12 Fig. 1.2 Electricity generation by fuel type in EU [1]

Fig. 3 .

 3 Fig. 3.23 Partial pressure of hydrogen along the membrane in the SMR configuration at 10 bar and 923 K. Figure taken from the work by Koch et al. [3].

Fig. 5 . 3 Fig. 5 . 4

 5354 Fig. 5.3 Scheme of the pressure change for the dynamic tests. Dynamic tests regarding the response of the system to the fuel flow rate changes were performed through two schemes, i.e. intervals of 50 l/min, and intervals of β5 l/min, both between 50 and 100 l/min, as shown in Fig. 5.4.

Fig. 5 . 5

 55 Fig. 5.5 Simulink model to simulate the reactor pressure and the pure hydrogen production rate

. 2 T(

 2 378 , thermometer error: ±0.32 , u(thermometer)= 873 K) = 378 , K-type error: ±0.8 , u(K-type)= 923 K) = 393 , K-type error: ±0.9 , u(K-type)= K, thermometer error: ±1.8 , u(thermometer)= . √ = 1 K T = 873 K, K-type error: ±4.5 , u(K-type)= . √ = 2.6 K T = 923 K, thermometer error: ±2 , u(thermometer)= √ = 1= 923 K, K-type error: ±4.9 , u(K-type)= ) = ( ) + ( ) + ( )

  solutionner cet aspect est un stockage chimique « Power To Gas ». Il est aussi possible d'utiliser des biocarburants, comme le bio-méthanol ou le bioéthanol, comme support : facilement stockable. La recherche consiste à proposer un moyen de conversion efficace. La voie catalytique, pour alimenter une pile à combustible, est considérée dans cette thèse. Les équations principales de conversion sont :

	C 2 H 5 OH → H 2 + CO + CH 4 Décomposition de l'éthanol	(1)
	CO + H 2 O ⇆ H 2 + CO 2	Equation du gaz à l'eau (Water Gas Shift WGS)	(2)
	CH 4 + 2H

2 O ⇆ 4H 2 + CO 2 Réaction du méthane (Methane Steam Reforming MSR) (3) Une équation globale de réaction du bioéthanol donne pour une mole d'éthanol 6 moles d'hydrogène : CH 3 CH 2 OH + 3H 2 O ↔ 2CO 2 + 6H 2 ( 298 K = 173.3 kJ/mol)

  Un plan d'expérience a été établi en fonction des limites opératoires du banc d'essai, et les résultats expérimentaux sont évalués en fonction de: • la sélectivité des quatre produits principaux du vapo-reformage (hydrogène H 2 , monoxyde de carbone CO, dioxyde de carbone CO 2 , méthane CH 4 ) hors vapeur d'eau H 2 O, • la fraction d'hydrogène (aval/amont membre),

		Température (K)	Pression (bar)	( l/min) Débit de carburant	Composition SC
	Plage de variation	873 -923	1 -12	50 -100	1.6, 2, 3
	Pas de variation pour le plan d'expérience	50	2	50	-
		Tableu 1 -Operating conditions
	Les paramètres opératoires du réacteur sont la pression, la température, le débit et la
	composition du carburant SC (Steam to Carbon Ratio), voir Tableau 1.

vapo-reformage de l'éthanol ESR en configuration CMR est décrit. Le vapo-reformage est réalisé sur un support catalytique Pd-Rh/CeO 2 dans un réacteur équipé d'une membrane MR à base de palladium et d'argent Pd-Ag. Le débit d'hydrogène avant et après membre est analysé en fonction des conditions opératoires du banc d'essai. Les résultats sont ensuite comparer à une configuration plus commune dite SMR (Staged Membrane Reactor).

Table 3 .

 3 1 Operating conditions

Table 3 .

 3 [START_REF] López | Hydrogen production from ethanol over Pd-Rh/CeO2 with a metallic membrane reactor[END_REF] The hydrogen flux at T=923 K for a membrane area of 30.4 cm 2

Table 4 .

 4 1 Wall temperature and the heat loss rate at 873 and 923 K

Table 4 .

 4

	2 NASA Polynomials [4]

Table 4 .

 4 3 Physical properties and standard chemical exergy of species

Table 5 .

 5 1 General models for calculation of the molar flow rate of the ESR products

Table 5 .

 5 [START_REF] López | Hydrogen production from ethanol over Pd-Rh/CeO2 with a metallic membrane reactor[END_REF] The specific models

Table 5 .

 5 3 Models for calculation of the hydrogen permeation rate

Table 5 .

 5 4 Experimental conditions for the dynamic experiments

Table 5 .

 5 [START_REF]Intergovernmental Panel on Climate Change IPCC[END_REF] Fitting parameters of the steady state model for the ESR products production rate model (eq. 5.5)

Table 5 .

 5 [START_REF] Arrhenius | On the Influence of Carbonic Acid in the Air upon the Temperature of the Ground[END_REF] Fitting parameters of the pre-exponential factor model (eq. 5.9)

Table 5 .

 5 [START_REF] Bernstein | Climate Change 2007: An Assessment of the Intergovernmental Panel on Climate Change[END_REF] Fitting parameters of the black box model (eq. 5.10)

	31

  The second appendix (appendix B) is devoted to the full results of the modeling of the molar flow rate of the ESR species (H 2 , CO, CO 2 , CH 4 , H 2 O) in the CMR. The formulation of the models and the parity plots are presented and are supported by tables of the fitting parameters for the species.

	the third chapter, the experimental results of the ethanol steam reforming (ESR) over the
	Pd-Rh/CeO 2 in a membrane reactor (MR) containing Pd-Ag membranes are given. The
	experimental results include the selectivity of the ESR products (H 2 , CO, CO 2 , and CH 4 ),
	hydrogen permeation rate (pure hydrogen production rate), hydrogen yield, and hydrogen
	recovery. The effect of the operating conditions (pressure, temperature, fuel flow rate, and
	the composition of the fuel, i.e. the molar ratio of water to ethanol) on the mentioned
	evaluation factors are discussed in details. Finally, a comparison between the reforming
	system presented in this thesis (CMR) and the Staged Membrane Reactor (SMR)
	configuration reported in the literature [2,3] is given.
	Chapter 4 is devoted to the comprehensive thermodynamic analysis of the CMR based on
	thermal and exergy efficiency using the experimental results. Exergy factors such as exergy
	destruction, unused exergy, and exergy efficiency are discussed and the results are
	presented via several figures and tables. The sources of exergy loss and the possible ways
	of the exergy recovery and optimization of the reformer in terms of the exergy efficiency
	are given. The advantages of the exergy analysis over the traditional thermodynamic
	analysis (based on thermal efficiency) are stated. This chapter also includes the
	thermodynamic analysis of the Staged Membrane Reactor (SMR) base on the experimental
	results reported in the literature for this configuration in ESR environment at similar
	operating conditions. A comparison between the CMR and the SMR also is presented.
	The results of the modeling and simulation of the reformer (the CMR) are presented in
	chapter 5. The methodology to develop the static models and dynamic simulation is
	discussed. Static models to calculate the molar production rate of ESR products, together
	with the calculation of the permeation rate of hydrogen through the membrane are given.
	Arrhenius' law based formulation is used to model the molar production rate of ESR
	products. In the case of the pure hydrogen production rate, two models, i.e. a black box
	model and a model based on the Sieverts' law are presented. The dynamic simulation
	model of the hydrogen permeation rate (transient state) is developed based on ideal gas law
	assumptions by using a first order function solved by Ordinary Differential Equation
	(O.D.E) solver.
	Two appendixes are given in this thesis. The first appendix (appendix A) is devoted to the
	analysis of the errors and uncertainty of the experimental measurements and exergy
	analysis. Appendix A is presented to support the data given via tables and figures and to
	demonstrate the range of the validity based on the maximum errors.

  Co supported on Al 2 O 3 and SiO 2 was tested by Batista et al.[START_REF] Batista | High efficiency steam reforming of ethanol by cobalt-based catalysts[END_REF] and complete conversion and selectivity of 65% were shown at 673 K. The concentration of CO in the gaseous mixture dropped to 800 ppm in the case of Co/Al 2 O 3 with 18% Cobalt. The authors reported that the produced CO reacted with water The catalysts were prepared using a Co precursor dissolved in methanol. The Co/SiO 2 showed the highest selectivity towards hydrogen. The carbon formation over the catalyst surface after the reaction was characterized by means of Raman spectroscopy. The Raman results of the samples after the catalytic tests showed that the type of carbon formed was affected by the particle size. Co-based catalysts over several supports such as ZnO, MgO, Al 2 O 3 , SiO 2 -TiO 2 , V 2 O 5 , La 2 O 3 , CeO 2 , and Sm 2 O 3 were prepared by Llorca et al.[START_REF] Llorca | Efficient Production of Hydrogen over Supported Cobalt Catalysts from Ethanol Steam Reforming[END_REF][START_REF] Llorca | In situ magnetic characterisation of supported cobalt catalysts under steam-reforming of ethanol[END_REF] and tested at 573-723 K in the ESR environment. The best catalytic activity was shown by Co/ZnO so that at 100% ethanol conversion, selectivity up to 73.8% to H 2 and 24.2% to CO 2 was obtained. After the reaction, Co particles were covered by disorder carbon. Carbon deposition was found on the surface of the ZnO support as well. In another study by the same authors[START_REF] Llorca | First use of macroporous silicon loaded with catalyst film for a chemical reaction: A microreformer for producing hydrogen from ethanol steam reforming[END_REF], they prepared Co catalysts by coating the micromonolithic support with Co 3 O 4 -ZnO. They reported more than 73% volume of hydrogen in the gaseous product at 773 K and S/C=3. Cobalt supported on CeO 2 has shown superior activity and hydrogen yield compared to other commercial supports of similar particle size at 673-773

(water gas shift reaction) or with hydrogen (methanation), which resulted in higher conversion over Co catalyst. Co/Al 2 O 3 showed superior activity in terms of CO removal. Lucredio et al.

[START_REF] Lucredio | Co catalysts supported on SiO2 and -Al2O3 applied to ethanol steam reforming: Effect of the solvent used in the catalyst preparation method[END_REF]

tested cobalt catalysts supported on Al 2 O 3 and SiO 2 in the ESR environment. K

[START_REF] Soykal | Ethanol steam reforming over Co/CeO2 catalysts: Investigation of the effect of ceria morphology[END_REF][START_REF] Bayram | Ethanol steam reforming over Co-based catalysts: Investigation of cobalt coordination environment under reaction conditions[END_REF][START_REF] Silva | The effect of support reducibility on the stability of Co/CeO2 for the oxidative steam reforming of ethanol[END_REF]

. The greater performance of Co/CeO 2 nanocubes is attributed to improved metal dispersion, increased reducibility, and higher oxygen mobility

[START_REF] Soykal | Ethanol steam reforming over Co/CeO2 catalysts: Investigation of the effect of ceria morphology[END_REF]

. Li et al.

[START_REF] Lin | Hydrogen Production from Ethanol Steam Reforming Over Supported Cobalt Catalysts[END_REF]

studied the synergic effect of ZrO 2 and CeO 2 (to suppress the methanation of the ZrO 2 ) on ethanol conversion promotion. The Co/CeZrO 2 showed higher activity than Co/ZrO 2 so that ethanol conversion of 100% and hydrogen yield of 82% (4.9 mol H 2 /mol ethanol) at 723 K was reached.

  2 , CeO 2 , Al 2 O 3 , Al 2 O 3 -La 2 O 3 , and ZrO 2 have been tested and in all cases improved activity, durability and resistance against coke formation has been reported[START_REF] Vinci | Ethanol Steam Reforming in a Two-Step Process . Short-Time Feasibility Tests[END_REF][START_REF] Zhang | Hydrogen Production via Steam Reforming of Ethanol on Phyllosilicate-Derived Ni/SiO β μ Enhanced εetal -Support Interaction and Catalytic Stability[END_REF][START_REF] Xu | Steam Reforming of Ethanol on Ni/CeO β μ Reaction Pathway and Interaction between Ni and the CeO 2 Support[END_REF][START_REF] Cerritos | Steam Reforming of Ethanol over Ni/Al2O3-La2O3 Catalysts Synthesized by Sol-Gel[END_REF][START_REF] Li | Steam reforming of ethanol over Ni/ZrO2 catalysts: Effect of support on product distribution[END_REF]. Ni/Al 2 O 3 catalyst mixed with CeO 2 was tested by Laosiripojana and Assabusrungrat[START_REF] Laosiripojana | Catalytic steam reforming of ethanol over high surface area CeO2: The role of CeO2 as an internal pre-reforming catalyst[END_REF]. The results showed that while the selectivity of the catalyst toward hydrogen was lower compared to pure Ni/Al 2 O 3 , the resistance against carbon formation was considerably higher. Chen et al.[START_REF] Chen | Regenerable and durable catalyst for hydrogen production from ethanol steam reforming[END_REF] used perovskite-type oxides La 1-x Ca x Fe 0.7 Ni 0.3 O 3 as an improved catalyst for ESR process capable to resist against carbon deposition. They found it promising to regulate the redox ability of the perovskite-type oxides for carbon deposition prevention. Many other additives such as copper, lanthanide, molybdenum, calcium, and magnesium in combination with nickel have been investigated and reported in the literature[START_REF] Llorca | Hydrogen from bio-ethanol[END_REF][START_REF] Youn | Hydrogen production by auto-thermal reforming of ethanol over Ni/ -Al2O3 catalysts: Effect of second metal addition[END_REF][START_REF] Youn | Role and effect of molybdenum on the performance of Ni-εo/ -Al2O3 catalysts in the hydrogen production by autothermal reforming of ethanol[END_REF][START_REF] Vizcaino | Ethanol steam reforming on Ni/Al2O3 catalysts: Effect of Mg addition[END_REF][START_REF] Özkan | Active carbon-supported Ni, Ni/Cu and Ni/Cu/Pd catalysed steam reforming of ethanol for the production of hydrogen[END_REF]. Adding elements such as La and Mg decreased the formation of C 2 H 4 and increased the activity of the catalyst and the selectivity towards hydrogen[START_REF] Sanchezsanchez | Ethanol steam reforming over Ni/La-Al2O3 catalysts: Influence of lanthanum loading[END_REF][START_REF] Vizcaino | Ethanol steam reforming on Ni/Al2O3 catalysts: Effect of Mg addition[END_REF]. Hernandez et al.[START_REF] Hernández | Steam reforming of ethanol over nickel-tungsten catalyst[END_REF] added W to Al 2 O 3 to stabilize the Ni-based catalyst. The surface area of the catalyst decreased and the pore volume of the mesoporous materials increased. Hydrogen selectivity of around 70% was reached at 823 K on alumina with Ni-W. Complete conversion and selectivity of 70% over NiAl 2 O

[START_REF]NASA polynomials[END_REF]

spinels was reported by Barroso et al

[START_REF] Barroso | Reactivity of aluminum spinels in the ethanol steam reforming reaction[END_REF]

. Ni/MgAl 2 O 4 was doped with Ce and in this case the selectivity was slightly improved and increased to 74% at 923 K after 10 h

[START_REF] Galetti | Promotion of Ni/MgAl2O4 catalysts with rare earths for the ethanol steam reforming reaction[END_REF]

. Biswas and Kunzru studied the adding of Cu and Ca to Ni supported on CeO 2 -ZrO 2 and reported an enhancement in catalyst activity and promotion of the WGSR

[START_REF] Biswas | Steam reforming of ethanol on Ni-CeO2-ZrO2 catalysts: Effect of doping with copper, cobalt and calcium[END_REF]

. Ca was the best dopant in terms of hydrogen production. ZrO 2 played the role of stabilizer in Ni-based catalysts.

  2 , ZrO 2 , and Al 2 O 3 are among the most used supports for the noble metals. The catalytic activity of Pt supported by Al 2 O 3 , ZrO 2 , and Al 2 O 3 -ZrO 2 was investigated by Domok et al. [88] in terms of the performance of catalyst in conversion of ethanol and CO. Ciambelli et al. [89] also tested Pt/CeO 2 catalyst for ESR at 573K and reported total conversion of ethanol and very low production of CO. CeO 2 -supported Pt/Ni catalyst for conversion of biomass derived ethanol was synthetized and tested by Palma et al.

  2 , CO 2 , CH 4 , CO, C 2 H 4 , C 2 H 6 , C 2 H 4 O and (C 2 H 5 ) 2 O was reported at 300-723 K.

More investigations on Rh can be found in the literature compared to Pt or Pd. For ethanol steam reforming reactions, Rh-based catalysts are very active and well known for the high selectivity towards hydrogen. Hydrogen selectivity of 73.5% at 873 K over Rh/Al 2 O 3

  0.8 Zr 0.2 O 2 . The best performance in terms of hydrogen yield was reported at 723 K in the case of a 2%Rh/Ce 0.8 Zr 0.2 O 2 catalyst, prepared via impregnation without pre-calcination of support. The high activity of this catalyst (100% ethanol conversion and 4.3 mol hydrogen per mol ethanol at 723 K) was ascribed to the strong interaction between Rh and Ce 0.8 Zr 0.2 O 2 and the high oxygen transfer rate favoring reforming of acetaldehyde instead of methane production. Rh-based catalysts were prepared by co-precipitation of aqueous solutions of Ce and Zr (and/or Pr and Y) nitrate and tested at 723 K and pressure up to 15 bar in the ESR environment[START_REF] Karim | A comparative study between Co and Rh for steam reforming of ethanol[END_REF]. According to the reaction pathway of ESR over Ru, the selectivity towards methane reaches an equilibrium leading to high selectivity at higher pressures. The iron oxides in the vicinity of Rh sites reduce the CO adsorption on Rh sites resulting in high hydrogen selectivity of 66% at 623 K. Iron oxide promoted the WGSR via which CO is converted to CO 2 . In another study, it was shown that addition of Ni to Rh-Y 2 O 3 -Al 2 O 3 wildly improved the catalytic reactivity correlated to increase in the methane steam reforming activity[START_REF] Valant | Preparation and characterization of bimetallic Rh-Ni/Y2O3-Al2O3 for hydrogen production by raw bioethanol steam reforming: influence of the addition of nickel on the catalyst performances and stability[END_REF].Rh supported on CeO 2 was tested by Da Silva et al.[START_REF] Da Silva | Steam and CO2 reforming of ethanol over Rh/CeO2 catalyst[END_REF]. They used both low and high surface area ceria and concluded that despite the deactivation of the catalysts, no carbon formed on the catalyst. Rh/CeO 2 /Al 2 O 3 catalyst at temperature range of 673-873 K was tested by Peela et al.[START_REF] Peela | Steam reforming of ethanol over Rh/CeO2/Al2O3 catalysts in a microchannel reactor[END_REF] for ESR and they found that when ceria was added to 2% Rh/Al 2 O 3 , the catalytic activity was improved. The test was done using two configurations of packed bed reactor and microchannel reactor. It was shown that although the hydrogen yield of microchannel reactor was slightly higher, the overall reactivity stayed the same for both configurations. The coke formation decreased by 3.5 times. Wanat et al. studied Rh and Rh-Ce catalyst in water gas shift and steam reforming of ethanol reactors[START_REF] Wanat | Steam reforming and water-gas shift of ethanol on Rh and Rh-Ce catalysts in a catalytic wall reactor[END_REF]. More than 99% conversion and methane selectivity of less than 1% were obtained with Rh-Ce catalyst at 1073 K. Very high hydrogen selectivity (98%) was obtained by coupling the steam reforming reactor with combustion Pt-based catalyst (Pt-Ce). Rh and Rh-Co catalysts supported on ZrO 2 , Al 2 O 3 , MgO, and Mg-Al showed high catalytic activity and carbon formation was lower specially on the bimetallic noble metal catalysts[START_REF] Moura | Ethanol steam reforming over rhodium and cobalt-based catalysts: Effect of the support[END_REF][START_REF] Churakova | Bimetallic Rh-Co/ZrO2 catalysts for ethanol steam reforming into hydrogen-containing gas[END_REF].Gucciardi et al[START_REF] Gucciardi | Ethanol and dimethyl ether steam reforming on Rh/Al2O3 catalysts for high-temperature fuel-cell feeds[END_REF] used Rh/Al 2 O 3 for both bio-ethanol and DME steam reforming at atmospheric pressure and temperature range of 823-923 K. They studied the catalytic activity and stated that the catalyst was less prone to coking in case of DME. Rh and Cobased catalysts supported on MgO, Al 2 O 3 , and Mg-Al oxides were evaluated for the ESR by Moura et al.[START_REF] Moura | Ethanol steam reforming over rhodium and cobalt-based catalysts: Effect of the support[END_REF]. In terms of the hydrogen yield, ethane selectivity, and specific surface area during the reaction, Mg-Al supported Rh and Co catalysts were introduced as the best ones. Lopez et al.[START_REF] López | Hydrogen production from ethanol over Pd-Rh/CeO2 with a metallic membrane reactor[END_REF] evaluated the performance of Pd-Rh/CeO They used CeO 2 , ZrO 2 , and Al 2 O 3 as the supports and concluded that Rh and Pt, when supported on CeO 2 /ZrO 2 showed the best activity of 100% ethanol conversion at around 923 K. Co, Ni, Rh, Rh-Co or Rh-Ni supported on mixed Zr-Ce oxides for redox characteristic enhancement were tested by Sarria et al.[START_REF] Romero-Sarria | Hydrogen production by steam reforming of ethanol[END_REF]. They mentioned that the stability of cobalt catalyst together with resistance against carbon formation was remarkably improved by addition of small amount of Rh. The enhanced activity of Ni/CeZr catalyst for ESR process was reported thanks to the addition of Rh. D.K. Liguras et al.[START_REF] Liguras | Production of hydrogen for fuel cells by steam reforming of ethanol over supported noble metal catalysts[END_REF] comprehensively studied the active metallic phases of Rh, Ru, Pt and Pd over different supports of Al 2 O

Huang et al. [101]

studied CeO 2 -supported catalysts for low temperature (573 K) ethanol steam reforming using a multi-channel reactor. The addition of Co to Rh/CeO 2 results in decreased catalytic selectivity towards CO and CH 4 . Catalyst deactivation was seen probably due to catalyst sintering, metal oxidation and coke deposition during ESR. Wu and Kawi

[START_REF] Kawi | Steam reforming of ethanol to H2 over Rh/Y2O3: crucial roles of Y2O3 oxidizing ability, space velocity, and H2/C[END_REF]

studied Ru-based catalysts based on several supports at 923-1073 K. Ru/Y 2 O 3 produced the highest hydrogen rate via ethanol steam reforming due to the synergic effect of electron-accepting Y 2 O 3 electron donating Rh. Thanks to the strong ability of Y 2 O 3 , the hydrogen of ethanol was oxidized to produce gaseous hydrogen and reduced the hydrogen from water. In general, the order of reactivity was Rh/Y 2 O 3 > Rh/ CeO 2 > Rh/La 2 O 3 > Rh/Al 2 O 3 . For the first time a novel iron promoted Rh-based catalyst (Rh-Fe/Ca-Al 2 O 3 ) was prepared by Chen et al.

[START_REF] Chen | Carbon monoxide-free hydrogen production via low-temperature steam reforming of ethanol over iron-promoted Rh catalyst[END_REF]

to produce a CO free hydrogen rich gas via ethanol steam reforming at low temperature (623-673 K). 2 for ESR by using a membrane reactor for online hydrogen purification. The effect of different operational conditions was studied and the best performance was obtained at 923 K and pressure of 9-11 bar. The same catalyst was tested by Divins et al.

[START_REF] Divins | Bio-ethanol steam reforming and autothermal reforming in 3-m channels coated with RhPd/CeOβ for hydrogen generation[END_REF]

. The microchannels of the silicon micromonolith were coated with Pt-Rh/CeO 2 and were tested for ESR process at 873 K using bio-ethanol with S/C of 2. Up to 3.8 moles hydrogen per one mole of ethanol was obtained and a significant improvement in the reactivity was cited compared to the conventional cordierite monoliths. In another work by Cobo et al.

[START_REF] Cobo | Steam reforming of ethanol over bimetallic RhPt/La2O3: Long-term stability under favorable reaction conditions[END_REF]

Pt-Rh bimetallic catalyst supported on La 2 O 3 was selected for ESR process. Temperature of 873 K together with the S/C of 7 was chosen for the experiments. 99% of the ethanol was converted and small carbon deposition was detected that was attributed to the formation of Rh-Pt-Rh 2 O 3 sites. Cavallaro

[START_REF] Cavallaro | Ethanol Steam Reforming on Rh/Al2O3 Catalysts[END_REF]

used Rh/Al 2 O 3 catalyst for ESR at 323-923 K and reported that at 923 K no carbon deposition was detected and the reactivity of the catalyst was maintained for several hours. Complete ethanol conversion was reached at S/C ratio of 4.2. A range of noble metals including Rh, Pd and Pt was evaluated by J.P. Breen et al.

[START_REF] Breen | Metal-catalysed steam reforming of ethanol in the production of hydrogen for fuel cell applications[END_REF]

for ESR process at 673-1023 K. 3 , MgO and TiO 2 for ESR process at 873-1123 K. They reported that Ru/Al 2 O 3 was able to fully convert ethanol with a hydrogen selectivity of 95%.

Table 3 .

 3 .2. 2 The hydrogen flux at T=923 K for a membrane area of 30.4 cm 2

	Partial pressure of hydrogen (atm)	20	12	7.2	3.4
	P ,	. -P ,	. (atm 0.5 )	3.5	2.5	1.7	0.9
	Hydrogen flux (mol/s)	2.6×10 -4	1.9×10 -4	1.4×10 -4	8.1×10 -5

Table 4 .

 4 .3 according to the Bejan model II[START_REF] Adrian Bejan | Thermal Design and optimization[END_REF]. 3 Physical properties and standard chemical exergy of species

	Component	Density (kg/m 3 )	Molar Weight (g/mol)	Standard chemical exergy (J/mol)
	H 2	0.0813	2.016	236100
	CH 4	0.648	16.04	831650
	CO	1.13	28.01	275100
	CO 2	1.775	44.01	19870
	Water	1000	18.02	900
	Ethanol	783	46.07	1357700

Table 5 .

 5 .5. 5 Fitting parameters of the steady state model for the ESR products production rate model (eq. 5.5)

	specie	α specie [mol.m -3 ]	specie [-]	specie [J.mol -1 .Pa -1 ]	specie [J.mol -1 ]	R 2
	H 2	1.1	0.71	8.4×10 -7	-0.1	0.995
	CO	75.5	0.9	-1.5×10 -6	-4.3	0.985
	CO 2	133.9	1.1	-1.5×10 -7	1.2	0.991
	CH 4	560.3	1.3	-4.8×10 -7	3.7	0.995
	H 2 O	226.1	1.1	-1.8×10 -7	2.7	0.999

Table 5 .

 5 2 ) increases from 0.68 to 0.91 when the measured value at 6 bar are not presented due to 7 Fitting parameters of the black box model (eq. 5.10)

	Parameter	a	b	c	e
	value	2.1×10 -11	0.7	1.5	3.1×10 -6
	Unit	[Pa -1 ]	[-]	[-]	[mol.s -1 ]

  3. A steady state model was proposed based on the Arrhenius law as a function of fuel flow rate and reactor pressure. The model was based on the experimental results and predicted correctly the performance of the ESR in terms of the molar production rate of ESR products, i.e. CO, CO 2 , CH 4 , H 2 , and H 2 O inside the reactor (catalytic zone). The pure hydrogen production rate (permeate zone) at steady state conditions was simulated by means of two steady state models, i.e. the Sieverts' law model named as model 1, and the black box model named as model 2.

The results of the uncertainty analysis of the exergy analysis

  All exergy factors are given in Watt[W].

		T [K] 923 Parameter T [K] 923 Parameter T [K] 923 Parameter T [K] 923 Parameter T [K] 923 Parameter T [K] 873 Parameter T [K] 873 Parameter T [K] 873 Parameter T [K] 873 Parameter T [K] 873 Parameter T [K] 873 Parameter 973 T [K] 973 T [K] 973 T [K] 973 T [K] 973 T [K] 973 T [K] 873 T [K] 873 T [K] 873 T [K] 873 T [K] 873 T [K] 873 A.4 T [K]	FF [μl/min] 100 FEtOH [mol/s] FF [μl/min] 50 FEtOH [mol/s] FF [μl/min] 100 FEtOH [mol/s] FF [μl/min] 50 FEtOH [mol/s] FF [μl/min] 100 FEtOH [mol/s] FF [μl/min] 50 FEtOH [mol/s] FF [μl/min] 100 FEtOH [mol/s] FF [μl/min] 50 FEtOH [mol/s] FF [μl/min] 100 FEtOH [mol/s] FF [μl/min] 50 FEtOH [mol/s] FF [μl/min] 100 FEtOH [mol/s] 50 FF [μl/min] FF [μl/min] 100 FF [μl/min] 50 FF [μl/min] 100 FF [μl/min] 50 FF [μl/min] 100 FF [μl/min] 50 FF [μl/min] 100 FF [μl/min] 50 FF [μl/min] 100 FF [μl/min] 50 FF [μl/min] 100 B.1 General Models S/C [-] FH2O [mol/s] S/C [-] FH2O [mol/s] S/C [-] FH2O [mol/s] S/C [-] FH2O [mol/s] S/C [-] FH2O [mol/s] S/C [-] FH2O [mol/s] S/C [-] FH2O [mol/s] S/C [-] FH2O [mol/s] S/C [-] FH2O [mol/s] S/C [-] FH2O [mol/s] S/C [-] FH2O [mol/s]	S/C 1.6 GC [%] S/C 2 GC [%] S/C 2 GC [%] S/C 3 GC [%] S/C 3 GC [%] S/C 1.6 GC [%] S/C 1.6 GC [%] S/C 2 GC [%] S/C 2 GC [%] S/C 3 GC [%] S/C 3 1.6 GC [%] S/C S/C 1.6 S/C 2 S/C 2 S/C 3 S/C 3 S/C 1.6 S/C 1.6 S/C 2 S/C 2 S/C 3 S/C 3
	T [K] Value Value Value Value Value Value Value Value Value Value Value Parameter Parameter Parameter Parameter Parameter Parameter Parameter Parameter Parameter Parameter Parameter 923 Parameter Value Uncertainty (u) Molar flow rate of the species present in the retentate stream (mol/s) FF [μl/min] 1.6 1.42E-05 2 6.33E-06 2 1.27E-05 3 4.93E-06 3 9.87E-06 1.6 7.08E-06 1.6 1.42E-05 2 6.33E-06 2 1.27E-05 3 4.93E-06 3 9.87E-06 EXH2.perm [W] EXfuel [W] EXH2.perm [W] EXfuel [W] EXH2.perm [W] EXfuel [W] EXH2.perm [W] EXfuel [W] EXH2.perm [W] EXfuel [W] EXH2.perm [W] EXfuel [W] EXH2.perm [W] EXfuel [W] EXH2.perm [W] EXfuel [W] EXH2.perm [W] EXfuel [W] EXH2.perm [W] EXfuel [W] EXH2.perm [W] EXfuel [W] 50 S/C [-] FEtOH [mol/s] 1.6 7.08E-06 0.014 7.10E-07 P [bar] CO2 H2 CH4 CO H2O CO2 4 3.63E-06 4.01E-06 7.21E-06 2.49E-07 1.65E-05 4.52E-07 6 3.88E-06 3.94E-06 6.72E-06 3.22E-07 1.56E-05 4.11E-07 8 3.61E-06 1.99E-06 5.03E-06 3.16E-07 1.22E-05 3.17E-07 10 2.55E-06 1.79E-06 4.20E-06 3.11E-07 1.06E-05 2.40E-07 12 2.00E-06 1.51E-06 3.30E-06 2.95E-07 8.77E-06 1.40E-07 P [bar] YH2 u(YH2) RH2 u(RH2) H2 12 0.20 0.022 0.79 0.138 0.13 12 0.15 0.013 0.68 0.026 0.10 12 0.48 1.08 0.26 0.40 1.48 1.74 3.01E-05 4.62E-05 2.56E-05 5.11E-05 3.01E-05 6.03E-05 2.31E-05 4.62E-05 2.56E-05 5.11E-05 6.03E-05 Tr [K] Tr [K] Tr [K] Tr [K] Tr [K] Tr [K] Tr [K] Tr [K] Tr [K] Tr [K] Tr [K] FH2O [mol/s] 2.31E-05 2.80E-06 Uncertainty (mol/s) 5 5 5 5 5 5 5 5 5 5 5 Tw [K] Tw [K] Tw [K] Tw [K] Tw [K] Tw [K] Tw [K] Tw [K] Tw [K] Tw [K] Tw [K] S/C 1.6 GC [%] 5 2.9 H2 CH4 4.52E-07 4.56E-07 4.51E-07 CO 4.11E-07 4.12E-07 4.10E-07 3.16E-07 3.19E-07 3.16E-07 2.38E-07 2.47E-07 2.36E-07 1.38E-07 1.45E-07 1.36E-07 u( H2) FH2.perm (mol/s) 0.014 5.78E-06 0.008 9.07E-06 0.10 Uncertainty (u) 0.014 7.10E-07 2.80E-06 2.9 Molar flow rate of the species present in the retentate stream (mol/s) Uncertainty (mol/s) P [bar] CO2 H2 CH4 CO H2O CO2 H2 CH4 CO 4 8.29E-06 9.52E-06 1.67E-05 4.43E-07 3.77E-05 1.05E-06 1.05E-06 1.07E-06 1.04E-06 6 7.51E-06 7.95E-06 1.39E-05 5.71E-07 3.22E-05 9.73E-07 9.73E-07 9.79E-07 9.72E-07 8 9.27E-06 4.84E-06 1.36E-05 4.36E-07 3.14E-05 8.24E-07 8.23E-07 8.28E-07 8.22E-07 10 8.38E-06 3.15E-06 1.05E-05 3.73E-07 2.51E-05 6.72E-07 6.67E-07 6.76E-07 6.67E-07 12 6.30E-06 2.54E-06 9.09E-06 3.21E-07 2.23E-05 5.52E-07 5.48E-07 5.58E-07 5.47E-07 P [bar] YH2 u(YH2) RH2 u(RH2) H2 u( H2) FH2.perm (mol/s) 4 0.04 0.007 0.27 0.039 0.03 0.007 3.54E-06 6 0.15 0.010 0.61 0.031 0.11 0.009 1.25E-05 8 0.22 0.013 0.79 0.028 0.19 0.010 1.84E-05 10 0.28 0.016 0.88 0.022 0.25 0.013 2.39E-05 12 0.35 0.019 0.92 0.016 0.31 0.015 3.00E-05 Uncertainty (u) 0.015 3.55E-07 1.40E-06 2.9 Molar flow rate of the species present in the retentate stream (mol/s) Uncertainty (mol/s) P [bar] CO2 H2 CH4 CO H2O CO2 H2 CH4 CO 4 3.88E-06 4.70E-06 6.64E-06 1.29E-07 2.00E-05 4.47E-07 4.47E-07 4.50E-07 4.45E-07 6 3.06E-06 3.32E-06 5.33E-06 8.95E-08 1.73E-05 3.46E-07 3.48E-07 3.66E-07 3.36E-07 8 3.75E-06 2.68E-06 4.69E-06 1.06E-07 1.61E-05 3.27E-07 3.24E-07 3.31E-07 3.21E-07 10 3.07E-06 1.95E-06 3.57E-06 8.17E-08 1.38E-05 2.52E-07 2.50E-07 2.54E-07 2.49E-07 12 3.06E-06 1.70E-06 3.34E-06 7.29E-08 1.33E-05 2.52E-07 2.42E-07 2.55E-07 2.37E-07 P [bar] YH2 u(YH2) RH2 u(RH2) H2 u( H2) FH2.perm (mol/s) 4 0.05 0.015 0.31 0.062 0.05 0.013 2.08E-06 6 0.18 0.018 0.67 0.030 0.15 0.014 6.72E-06 8 0.29 0.022 0.81 0.021 0.25 0.016 1.11E-05 10 0.39 0.027 0.88 0.014 0.33 0.018 1.48E-05 12 0.47 0.030 0.91 0.012 0.40 0.020 1.78E-05 Uncertainty (u) 0.015 7.10E-07 2.80E-06 2.9 Molar flow rate of the species present in the retentate stream (mol/s) Uncertainty (mol/s) P [bar] CO2 H2 CH4 CO H2O CO2 H2 CH4 CO 4 7.70E-06 9.04E-06 1.34E-05 2.38E-07 4.02E-05 5.22E-07 5.22E-07 5.24E-07 5.22E-07 6 7.58E-06 6.98E-06 1.21E-05 1.91E-07 3.74E-05 5.91E-07 5.91E-07 5.94E-07 5.90E-07 8 8.36E-06 5.32E-06 1.02E-05 2.14E-07 3.37E-05 6.39E-07 6.38E-07 6.41E-07 6.37E-07 10 8.56E-06 4.31E-06 9.01E-06 1.97E-07 3.13E-05 7.11E-07 7.10E-07 7.13E-07 7.10E-07 12 9.34E-06 4.38E-06 9.69E-06 2.10E-07 3.27E-05 6.37E-07 6.33E-07 6.38E-07 6.33E-07 P [bar] YH2 u(YH2) RH2 u(RH2) H2 u( H2) FH2.perm (mol/s) 4 0.04 0.008 0.23 0.039 0.03 0.007 2.69E-06 6 0.13 0.011 0.59 0.025 0.11 0.008 1.01E-05 8 0.23 0.015 0.76 0.023 0.19 0.010 1.71E-05 10 0.30 0.018 0.84 0.022 0.25 0.012 2.25E-05 12 0.36 0.022 0.86 0.017 0.31 0.014 2.76E-05 Uncertainty (u) 0.02 2.77E-07 1.60E-06 2.9 Molar flow rate of the species present in the retentate stream (mol/s) Uncertainty (mol/s) P [bar] CO2 H2 CH4 CO H2O CO2 H2 CH4 CO 4 4.96E-06 6.10E-06 5.58E-06 2.05E-07 2.46E-05 5.16E-07 5.40E-07 5.29E-07 4.66E-07 6 3.37E-06 3.96E-06 3.46E-06 1.10E-07 3.03E-05 3.19E-07 3.25E-07 3.20E-07 3.02E-07 8 2.98E-06 2.76E-06 3.32E-06 7.18E-08 3.10E-05 2.68E-07 2.68E-07 2.69E-07 2.67E-07 10 3.47E-06 2.19E-06 2.90E-06 5.46E-08 2.98E-05 2.55E-07 2.50E-07 2.53E-07 2.46E-07 12 1.81E-06 1.81E-06 3.07E-06 3.06E-08 3.36E-05 2.14E-07 2.14E-07 2.24E-07 2.08E-07 P [bar] YH2 u(YH2) RH2 u(RH2) H2 u( H2) FH2.perm (mol/s) 4 0.02 0.020 0.07 0.082 0.01 0.013 4.70E-07 6 0.17 0.022 0.56 0.035 0.11 0.014 4.97E-06 8 0.34 0.027 0.79 0.019 0.23 0.016 1.01E-05 10 0.46 0.032 0.86 0.014 0.31 0.018 1.37E-05 12 0.53 0.036 0.90 0.011 0.35 0.019 1.58E-05 Uncertainty (u) 0.02 5.53E-07 3.21E-06 2.9 Molar flow rate of the species present in the retentate stream (mol/s) Uncertainty (mol/s) P [bar] CO2 H2 CH4 CO H2O CO2 H2 CH4 CO 4 1.01E-05 1.06E-05 1.16E-05 3.34E-07 4.88E-05 1.94E-06 2.03E-06 2.19E-06 7.51E-07 6 9.50E-06 8.07E-06 9.08E-06 2.76E-07 5.18E-05 7.47E-07 7.40E-07 7.46E-07 7.23E-07 8 6.26E-06 4.60E-06 6.28E-06 1.13E-07 6.21E-05 4.97E-07 4.86E-07 4.98E-07 4.74E-07 10 9.17E-06 6.04E-06 7.86E-06 1.57E-07 5.28E-05 6.48E-07 6.43E-07 6.47E-07 6.39E-07 12 5.18E-06 4.96E-06 8.47E-06 8.64E-08 6.18E-05 5.91E-07 5.90E-07 6.12E-07 5.78E-07 P [bar] YH2 u(YH2) RH2 u(RH2) H2 u( H2) FH2.perm (mol/s) 4 0.01 0.010 0.07 0.049 0.01 0.006 8.07E-07 6 0.17 0.014 0.55 0.027 0.11 0.008 9.88E-06 8 0.28 0.018 0.78 0.019 0.18 0.010 1.65E-05 10 0.37 0.023 0.78 0.019 0.24 0.012 2.16E-05 12 0.43 0.026 0.84 0.017 0.28 0.013 2.52E-05 Uncertainty (u) 0.014 7.10E-07 2.80E-06 2.9 Molar flow rate of the species present in the retentate stream (mol/s) Uncertainty (mol/s) P [bar] CO2 H2 CH4 CO H2O CO2 H2 CH4 CO 4 4.20E-06 3.53E-06 8.61E-06 2.21E-07 1.93E-05 5.40E-07 5.40E-07 5.42E-07 5.39E-07 6 3.44E-06 2.97E-06 7.49E-06 2.42E-07 1.71E-05 4.03E-07 4.03E-07 4.09E-07 4.02E-07 8 3.22E-06 2.35E-06 6.58E-06 2.12E-07 1.53E-05 3.47E-07 3.47E-07 3.50E-07 3.46E-07 10 3.05E-06 1.40E-06 5.34E-06 1.37E-07 1.27E-05 2.93E-07 2.93E-07 2.95E-07 2.93E-07 12 2.60E-06 1.03E-06 4.47E-06 1.30E-07 1.09E-05 2.48E-07 2.47E-07 2.50E-07 2.47E-07 P [bar] YH2 u(YH2) RH2 u(RH2) H2 u( H2) FH2.perm (mol/s) 4 0.02 0.014 0.16 0.117 0.03 0.013 6.69E-07 6 0.09 0.016 0.56 0.050 0.11 0.015 3.78E-06 8 0.13 0.019 0.71 0.037 0.19 0.016 5.65E-06 10 0.22 0.026 0.87 0.025 0.25 0.021 9.20E-06 12 0.28 0.031 0.92 0.018 0.31 0.025 1.19E-05 Uncertainty (u) 0.014 7.10E-07 2.80E-06 2.9 Molar flow rate of the species present in the retentate stream (mol/s) Uncertainty (mol/s) P [bar] CO2 H2 CH4 CO H2O CO2 H2 CH4 CO 4 8.85E-06 7.22E-06 1.87E-05 2.70E-07 4.14E-05 1.07E-06 1.07E-06 1.07E-06 1.07E-06 6 8.65E-06 6.07E-06 1.80E-05 4.63E-07 4.03E-05 9.47E-07 9.46E-07 9.53E-07 9.46E-07 8 8.70E-06 4.99E-06 1.71E-05 4.37E-07 3.83E-05 8.87E-07 8.85E-07 8.94E-07 8.85E-07 10 7.39E-06 3.85E-06 1.47E-05 1.78E-07 3.33E-05 7.62E-07 7.61E-07 7.66E-07 7.61E-07 12 6.80E-06 3.13E-06 1.36E-05 1.60E-07 3.12E-05 6.93E-07 6.93E-07 6.97E-07 6.93E-07 P [bar] YH2 u(YH2) RH2 u(RH2) H2 u( H2) FH2.perm (mol/s) 4 0.01 0.007 0.08 0.068 0.03 0.007 6.50E-07 6 0.05 0.007 0.43 0.049 0.11 0.007 4.51E-06 8 0.08 0.008 0.57 0.048 0.19 0.007 6.65E-06 10 0.14 0.010 0.75 0.038 0.25 0.008 1.16E-05 12 0.18 0.011 0.83 0.032 0.31 0.009 1.49E-05 Uncertainty (u) 0.015 3.55E-07 1.40E-06 2.9 Molar flow rate of the species present in the retentate stream (mol/s) Uncertainty (mol/s) P [bar] CO2 H2 CH4 CO H2O CO2 H2 CH4 CO 4 3.90E-06 3.96E-06 8.05E-06 8.80E-08 2.27E-05 9.77E-07 9.77E-07 9.82E-07 9.76E-07 6 3.84E-06 3.23E-06 8.27E-06 6.13E-08 2.32E-05 1.02E-06 1.02E-06 1.03E-06 1.02E-06 8 3.30E-06 2.69E-06 7.41E-06 4.92E-08 2.14E-05 1.17E-06 1.17E-06 1.18E-06 1.17E-06 10 2.64E-06 2.03E-06 6.05E-06 3.64E-08 1.87E-05 1.53E-06 1.52E-06 1.68E-06 1.50E-06 12 2.03E-06 1.43E-06 4.71E-06 2.78E-08 1.60E-05 1.98E-06 1.97E-06 2.06E-06 1.96E-06 P [bar] YH2 u(YH2) RH2 u(RH2) H2 u( H2) FH2.perm (mol/s) 4 0.02 0.015 0.18 0.104 0.02 0.013 8.74E-07 6 0.06 0.016 0.41 0.099 0.05 0.013 2.22E-06 8 0.09 0.016 0.56 0.116 0.08 0.013 3.36E-06 10 0.16 0.018 0.75 0.142 0.14 0.014 6.05E-06 12 0.23 0.020 0.86 0.167 0.20 0.015 8.74E-06 Uncertainty (u) 0.015 7.10E-07 2.80E-06 2.9 Molar flow rate of the species present in the retentate stream (mol/s) Uncertainty (mol/s) P [bar] CO2 H2 CH4 CO H2O CO2 H2 CH4 CO 4 8.31E-06 8.19E-06 1.72E-05 1.80E-07 4.77E-05 4.73E-07 4.73E-07 4.75E-07 4.73E-07 6 7.92E-06 6.72E-06 1.71E-05 1.32E-07 4.75E-05 4.94E-07 4.94E-07 4.98E-07 4.93E-07 8 7.47E-06 5.58E-06 1.60E-05 1.15E-07 4.53E-05 5.45E-07 5.43E-07 5.60E-07 5.41E-07 10 7.08E-06 4.83E-06 1.50E-05 9.75E-08 4.32E-05 5.88E-07 5.87E-07 5.91E-07 5.87E-07 12 6.62E-06 3.86E-06 1.33E-05 8.46E-08 3.97E-05 6.75E-07 6.73E-07 6.85E-07 6.72E-07 P [bar] YH2 u(YH2) RH2 u(RH2) H2 u( H2) FH2.perm (mol/s) 4 0.01 0.008 0.10 0.058 0.01 0.006 8.74E-07 6 0.04 0.008 0.29 0.046 0.03 0.007 2.69E-06 8 0.06 0.008 0.44 0.040 0.05 0.007 4.37E-06 10 0.10 0.009 0.62 0.034 0.09 0.007 7.73E-06 12 0.16 0.012 0.76 0.033 0.14 0.008 1.21E-05 Uncertainty (u) 0.02 2.77E-07 1.60E-06 2.9 Molar flow rate of the species present in the retentate stream (mol/s) Uncertainty (mol/s) P [bar] CO2 H2 CH4 CO H2O CO2 H2 CH4 CO 4 3.34E-06 4.13E-06 6.24E-06 5.32E-08 2.79E-05 5.79E-07 5.81E-07 5.88E-07 5.76E-07 6 2.95E-06 3.36E-06 5.60E-06 4.27E-08 2.66E-05 1.75E-06 1.96E-06 3.15E-06 6.56E-07 8 2.77E-06 2.89E-06 5.10E-06 3.89E-08 2.56E-05 7.42E-07 7.44E-07 7.93E-07 7.20E-07 10 2.13E-06 2.27E-06 3.90E-06 3.09E-08 2.32E-05 9.49E-07 9.49E-07 9.61E-07 9.45E-07 12 1.77E-06 1.58E-06 3.09E-06 2.52E-08 2.15E-05 1.30E-06 1.28E-06 1.43E-06 1.23E-06 P [bar] YH2 u(YH2) RH2 u(RH2) H2 u( H2) FH2.perm (mol/s) 4 0.01 0.019 0.05 0.127 0.00 0.013 2.02E-07 6 0.03 0.020 0.21 0.144 0.02 0.013 8.74E-07 8 0.04 0.020 0.37 0.113 0.04 0.013 1.21E-06 10 0.13 0.021 0.62 0.105 0.08 0.013 3.76E-06 Uncertainty (u) 0.02 5.54E-07 3.22E-06 2.9 Molar flow rate of the species present in the retentate stream (mol/s) Uncertainty (mol/s) P [bar] CO2 H2 CH4 CO H2O CO2 H2 CH4 CO 4 6.66E-06 8.30E-06 1.24E-05 1.12E-07 5.57E-05 2.88E-07 2.88E-07 2.89E-07 2.88E-07 6 6.35E-06 7.13E-06 1.21E-05 9.38E-08 5.50E-05 3.14E-07 3.15E-07 3.26E-07 3.09E-07 8 6.13E-06 5.96E-06 1.13E-05 8.21E-08 5.34E-05 3.36E-07 3.35E-07 3.47E-07 3.31E-07 10 5.98E-06 5.15E-06 1.05E-05 7.13E-08 5.16E-05 3.69E-07 3.69E-07 3.76E-07 3.66E-07 12 5.67E-06 4.27E-06 9.18E-06 7.04E-08 4.91E-05 4.27E-07 4.20E-07 4.53E-07 4.11E-07 P [bar] YH2 u(YH2) RH2 u(RH2) H2 u( H2) FH2.perm (mol/s) 4 0.00 0.010 0.02 0.066 0.00 0.006 2.02E-07 6 0.02 0.010 0.14 0.060 0.01 0.006 1.14E-06 8 0.04 0.010 0.28 0.051 0.03 0.007 2.35E-06 10 0.09 0.011 0.50 0.033 0.06 0.007 5.11E-06 10 0.50 1.09 0.41 0.55 1.64 2.05 0.11 8 0.51 1.10 0.52 0.66 1.76 2.28 0.12 12 0.67 1.76 0.90 1.05 2.81 3.71 0.20 0.019 5.85E-06 H2O 5.76E-06 5.57E-06 5.43E-06 5.23E-06 5.77E-07 5.77E-07 0.009 u(FH2.perm) H2O 7.06E-06 6.89E-06 6.59E-06 6.29E-06 6.05E-06 u(FH2.perm) 5.77E-07 5.77E-07 5.77E-07 5.77E-07 5.77E-07 H2O 3.37E-06 3.20E-06 3.13E-06 2.97E-06 2.98E-06 u(FH2.perm) 5.77E-07 5.77E-07 5.77E-07 5.77E-07 5.77E-07 H2O 5.98E-06 6.12E-06 6.22E-06 6.36E-06 6.21E-06 u(FH2.perm) 5.77E-07 5.77E-07 5.77E-07 5.77E-07 5.77E-07 H2O 3.49E-06 3.07E-06 2.97E-06 2.94E-06 2.88E-06 u(FH2.perm) 5.77E-07 5.77E-07 5.77E-07 5.77E-07 5.77E-07 H2O 9.25E-06 6.36E-06 5.86E-06 6.16E-06 6.09E-06 u(FH2.perm) 5.77E-07 5.77E-07 5.77E-07 5.77E-07 5.77E-07 H2O 6.02E-06 5.75E-06 5.63E-06 5.52E-06 5.43E-06 u(FH2.perm) 5.77E-07 5.77E-07 5.77E-07 5.77E-07 5.77E-07 H2O 7.08E-06 6.84E-06 6.72E-06 6.47E-06 6.33E-06 u(FH2.perm) 5.77E-07 5.77E-07 5.77E-07 5.77E-07 5.77E-07 H2O 1.96E-06 2.05E-06 2.35E-06 3.18E-06 4.02E-06 u(FH2.perm) 5.77E-07 5.77E-07 5.77E-07 5.77E-07 5.77E-07 H2O 9.48E-07 9.91E-07 1.10E-06 1.18E-06 1.36E-06 u(FH2.perm) 5.77E-07 5.77E-07 5.77E-07 5.77E-07 5.77E-07 H2O 3.61E-06 8.73E-06 4.02E-06 4.36E-06 5.30E-06 u(FH2.perm) 5.77E-07 5.77E-07 5.77E-07 5.77E-07 H2O 5.45E-06 5.52E-06 5.56E-06 5.62E-06 5.77E-06 u(FH2.perm) 5.77E-07 0.011 5.77E-07 12 0.48 0.97 0.41 0.55 1.52 1.94 0.10 0.010 12 0.67 1.66 1.02 1.17 2.83 3.85 0.20 0.020 12 0.45 0.84 0.37 0.51 1.35 1.72 0.09 0.009 12 0.63 1.41 0.97 1.11 2.53 3.50 0.18 0.018 12 0.47 1.07 0.42 0.56 1.62 2.04 0.11 0.011 12 0.67 1.76 1.11 1.25 3.01 4.12 0.22 0.021 12 0.74 1.24 2.96 3.10 4.34 7.31 0.37 0.036 12 0.65 1.64 1.08 1.22 2.86 3.93 0.20 0.020 12 0.62 1.01 2.01 2.15 3.17 5.18 0.26 0.025 12 0.58 1.36 0.73 0.87 2.23 2.95 0.16 0.015 0.012 5.77E-07 6 0.52 1.12 0.65 0.79 1.91 2.57 0.14 10 0.69 1.78 1.08 1.22 3.00 4.07 0.21 0.021 10 0.48 0.97 0.42 0.56 1.53 1.94 0.10 0.010 10 0.68 1.67 1.13 1.27 2.95 4.08 0.21 0.021 10 0.46 0.85 0.41 0.56 1.40 1.82 0.10 0.010 10 0.64 1.42 1.03 1.18 2.59 3.63 0.19 0.019 10 0.48 1.07 0.48 0.62 1.70 2.18 0.12 0.011 10 0.68 1.77 1.21 1.35 3.12 4.33 0.23 0.022 10 0.68 1.18 2.37 2.51 3.69 6.06 0.30 0.030 10 0.63 1.62 0.94 1.09 2.71 3.65 0.19 0.019 10 0.55 0.94 1.42 1.56 2.50 3.93 0.20 0.020 10 0.56 1.35 0.63 0.77 2.12 2.75 0.15 0.014 0.014 5.77E-07 Parameter EXH2.perm [W] EXfuel [W] Tr [K] Tw [K] Value Given in next table 9.6 923 393 Uncertainty (u) 0.14 0.6 4 0.73 P [bar] EXWel EXin EXretentate EXout EXdestruction EXunused EXH2.perm ex [%] ex [%] 4 12.1 21.7 7.5 8.3 12.4 20.0 0.7 3 7 6 12.2 21.8 7.1 9.2 11.6 18.7 2.1 10 19 8 12.5 22.1 5.1 8.2 12.9 18.0 3.0 14 28 10 12.6 22.2 4.3 8.3 12.9 17.2 3.9 19 35 12 12.7 22.3 3.4 8.2 13.1 16.6 4.8 22 42 P [bar] u(EXWel) u(Exin) u(EXretentate) u(EXout) u(EXdestruction) u(EXunused) u( ex) Non-Insulate reactor Insulate reactor 4 0.53 1.13 0.71 0.86 1.98 2.70 0.15 0.015 Value Given in next table 19.2 923 393 Uncertainty (u) 0.14 1.09 4 0.73 P [bar] EXWel EXin EXretentate EXout EXdestruction EXunused EXH2.perm ex [%] ex [%] 4 12.8 32.0 17.5 18.3 12.7 30.1 0.9 3 4 6 13.3 32.5 14.6 17.7 13.8 28.4 3.1 10 14 8 13.3 32.5 13.6 18.1 13.4 27.0 4.5 14 21 10 13.8 33.0 10.4 16.3 15.7 26.2 5.9 18 27 12 14.1 33.3 9.0 16.3 15.9 24.9 7.4 23 33 P [bar] u(EXWel) u(Exin) u(EXretentate) u(EXout) u(EXdestruction) u(EXunused) u( ex) Non-Insulate reactor Insulate reactor 4 0.75 1.84 1.63 1.78 3.62 5.25 0.28 0.029 6 0.74 1.83 1.52 1.66 3.49 5.01 0.27 0.027 8 0.72 1.81 1.30 1.44 3.25 4.54 0.24 0.024 Value Given in next table 8.6 923 393 Uncertainty (u) 0.14 0.49 4 0.73 P [bar] EXWel EXin EXretentate EXout EXdestruction EXunused EXH2.perm ex [%] ex [%] 4 12.1 20.6 7.2 7.8 11.9 19.1 0.5 3 5 6 12.3 20.9 5.7 7.4 12.5 18.2 1.6 8 17 8 12.4 21.0 5.0 7.8 12.2 17.3 2.7 14 27 10 12.6 21.1 3.8 7.5 12.7 16.5 3.6 18 36 12 12.6 21.2 3.6 7.9 12.2 15.8 4.4 22 43 P [bar] u(EXWel) u(Exin) u(EXretentate) u(EXout) u(EXdestruction) u(EXunused) u( ex) Non-Insulate reactor Insulate reactor 4 0.51 1.00 0.70 0.84 1.84 2.55 0.14 0.014 6 0.49 0.99 0.57 0.71 1.70 2.26 0.12 0.012 8 0.49 0.98 0.53 0.67 1.65 2.18 0.12 0.012 Value Given in next table 17.2 923 393 Uncertainty (u) 0.14 0.99 4 0.73 P [bar] EXWel EXin EXretentate EXout EXdestruction EXunused EXH2.perm ex [%] ex [%] 4 13.1 30.3 14.6 15.2 14.1 28.6 0.7 2 3 6 13.3 30.5 12.8 15.3 14.2 27.0 2.5 8 13 8 13.6 30.8 10.8 15.0 14.8 25.6 4.2 14 21 10 13.8 31.0 9.5 15.0 15.0 24.5 5.5 18 28 12 13.7 30.9 10.2 16.9 13.0 23.1 6.7 23 34 P [bar] u(EXWel) u(Exin) u(EXretentate) u(EXout) u(EXdestruction) u(EXunused) u( ex) Non-Insulate reactor Insulate reactor 4 0.65 1.64 0.86 1.00 2.64 3.50 0.20 0.020 6 0.66 1.65 0.96 1.10 2.75 3.72 0.20 0.020 8 0.67 1.66 1.03 1.17 2.83 3.86 0.21 0.021 Value Given in next table 6.7 923 393 Uncertainty (u) 0.14 0.39 4 0.73 P [bar] EXWel EXin EXretentate EXout EXdestruction EXunused EXH2.perm ex [%] ex [%] 4 12.0 18.7 6.8 6.9 10.7 17.6 0.1 1 1 6 12.3 19.0 4.6 5.8 12.3 16.8 1.2 7 15 8 12.4 19.1 4.2 6.6 11.4 15.6 2.5 14 31 10 12.4 19.1 3.7 7.0 11.1 14.8 3.4 19 41 12 12.4 19.1 3.7 7.6 10.5 14.2 3.9 21 48 P [bar] u(EXWel) u(Exin) u(EXretentate) u(EXout) u(EXdestruction) u(EXunused) u( ex) Non-Insulate reactor Insulate reactor 4 0.50 0.89 0.80 0.94 1.84 2.64 0.14 0.015 6 0.47 0.86 0.51 0.65 1.51 2.02 0.11 0.011 8 0.46 0.85 0.44 0.58 1.43 1.87 0.10 0.010 Value Given in next table 13.4 923 393 Uncertainty (u) 0.14 0.78 4 0.73 P [bar] EXWel EXin EXretentate EXout EXdestruction EXunused EXH2.perm ex [%] ex [%] 4 12.9 26.3 13.6 13.8 11.5 25.1 0.2 1 1 6 13.3 26.7 10.9 13.3 12.4 23.3 2.4 9 15 8 13.8 27.2 7.8 11.8 14.3 22.1 4.0 15 25 10 13.5 26.9 9.4 14.7 11.2 20.6 5.3 20 33 12 13.4 26.8 9.7 15.8 10.0 19.6 6.2 24 39 P [bar] u(EXWel) u(Exin) u(EXretentate) u(EXout) u(EXdestruction) u(EXunused) u( ex) Non-Insulate reactor Insulate reactor 4 0.89 1.67 2.82 2.96 4.63 7.45 0.39 0.040 6 0.65 1.43 1.17 1.32 2.75 3.92 0.21 0.021 8 0.61 1.39 0.81 0.95 2.35 3.16 0.17 0.017 Value Given in next table 9.6 873 378 Uncertainty (u) 0.14 0.6 3.6 0.7 P [bar] EXWel EXin EXretentate EXout EXdestruction EXunused EXH2.perm ex [%] ex [%] 4 11.8 21.4 8.6 8.8 11.6 20.2 0.2 1 2 6 12.0 21.6 7.5 8.4 12.2 19.6 0.9 4 9 8 12.1 21.7 6.5 7.9 12.8 19.3 1.4 7 13 10 12.3 21.9 5.2 7.4 13.5 18.7 2.2 11 21 12 12.4 22.0 4.3 7.2 13.8 18.1 2.9 14 26 P [bar] u(EXWel) u(Exin) u(EXretentate) u(EXout) u(EXdestruction) u(EXunused) u( ex) Non-Insulate reactor Insulate reactor 4 0.52 1.11 0.84 0.98 2.10 2.94 0.16 0.017 6 0.49 1.09 0.64 0.79 1.88 2.52 0.14 0.014 8 0.48 1.08 0.56 0.70 1.78 2.35 0.13 0.013 Value Given in next table 19.2 873 378 Uncertainty (u) 0.14 1.09 3.6 0.7 P [bar] EXWel EXin EXretentate EXout EXdestruction EXunused EXH2.perm ex [%] ex [%] 4 12.3 31.5 18.5 18.7 11.8 30.3 0.2 1 1 6 12.4 31.6 17.7 18.8 11.8 29.5 1.1 4 5 8 12.6 31.8 16.6 18.2 12.6 29.1 1.6 5 8 10 13.0 32.1 14.1 16.9 14.2 28.3 2.8 9 13 12 13.1 32.3 13.0 16.6 14.7 27.7 3.6 12 17 P [bar] u(EXWel) u(Exin) u(EXretentate) u(EXout) u(EXdestruction) u(EXunused) u( ex) Non-Insulate reactor Insulate reactor 4 0.73 1.82 1.65 1.79 3.61 5.27 0.28 0.029 6 0.71 1.80 1.48 1.62 3.42 4.90 0.26 0.026 8 0.70 1.79 1.39 1.53 3.32 4.72 0.25 0.025 Value Given in next table 8.6 873 378 Uncertainty (u) 0.14 0.49 3.6 0.7 P [bar] EXWel EXin EXretentate EXout EXdestruction EXunused EXH2.perm ex [%] ex [%] 4 11.7 20.3 8.2 8.5 10.9 19.1 0.2 1 2 6 11.7 20.3 8.3 8.8 10.5 18.7 0.5 3 6 8 11.8 20.4 7.4 8.2 11.2 18.6 0.8 4 9 10 12.1 20.6 6.0 7.5 12.2 18.2 1.5 8 15 12 12.3 20.9 4.7 6.8 13.1 17.7 2.1 11 22 P [bar] u(EXWel) u(Exin) u(EXretentate) u(EXout) u(EXdestruction) u(EXunused) u( ex) Non-Insulate reactor Insulate reactor 4 0.57 1.06 1.46 1.60 2.66 4.12 0.22 0.022 6 0.57 1.07 1.53 1.67 2.74 4.26 0.22 0.022 8 0.60 1.09 1.74 1.89 2.98 4.72 0.24 0.024 Value Given in next table 17.2 873 378 Uncertainty (u) 0.14 0.99 3.6 0.7 P [bar] EXWel EXin EXretentate EXout EXdestruction EXunused EXH2.perm ex [%] ex [%] 4 12.3 29.5 17.5 17.7 10.7 28.2 0.2 1 1 6 12.3 29.5 17.1 17.7 10.7 27.8 0.7 2 4 8 12.5 29.6 15.9 16.9 11.7 27.6 1.1 4 6 10 12.6 29.8 14.7 16.6 12.2 26.9 1.9 7 10 12 12.9 30.1 13.0 15.9 13.2 26.1 2.9 10 15 P [bar] u(EXWel) u(Exin) u(EXretentate) u(EXout) u(EXdestruction) u(EXunused) u( ex) Non-Insulate reactor Insulate reactor 4 0.61 1.61 0.78 0.92 2.52 3.30 0.19 0.019 6 0.62 1.61 0.81 0.95 2.56 3.37 0.19 0.019 8 0.63 1.62 0.89 1.03 2.65 3.54 0.19 0.019 Value Given in next table 6.7 873 378 Uncertainty (u) 0.14 0.39 3.6 0.7 P [bar] EXWel EXin EXretentate EXout EXdestruction EXunused EXH2.perm ex [%] ex [%] 4 11.8 18.5 6.8 6.9 10.6 17.4 0.0 0 1 6 11.9 18.6 6.1 6.3 11.3 17.4 0.2 1 3 8 12.0 18.7 5.5 5.8 11.9 17.4 0.3 2 4 10 12.2 18.9 4.3 5.2 12.7 16.9 0.9 5 12 12 12.3 19.0 3.4 4.8 13.2 16.6 1.4 8 18 P [bar] u(EXWel) u(Exin) u(EXretentate) u(EXout) u(EXdestruction) u(EXunused) u( ex) Non-Insulate reactor Insulate reactor 4 0.49 0.88 0.89 1.03 1.91 2.79 0.15 0.015 6 0.91 1.30 3.55 3.69 4.99 8.53 0.43 0.044 8 0.52 0.91 1.16 1.30 2.21 3.36 0.17 0.017 Value Given in next table 13.4 873 378 Uncertainty (u) 0.14 0.78 3.6 0.7 P [bar] EXWel EXin EXretentate EXout EXdestruction EXunused EXH2.perm ex [%] ex [%] 4 12.6 26.0 13.6 13.7 11.3 24.9 0.0 0 0 6 12.6 26.0 13.1 13.3 11.7 24.7 0.3 1 2 8 12.8 26.2 12.1 12.6 12.5 24.6 0.6 2 4 10 12.9 26.3 11.1 12.4 12.9 24.1 1.2 5 8 12 13.1 26.5 9.8 12.0 13.5 23.3 2.2 9 14 P [bar] u(EXWel) u(Exin) u(EXretentate) u(EXout) u(EXdestruction) u(EXunused) u( ex) Non-Insulate reactor Insulate reactor 4 0.55 1.33 0.51 0.65 1.98 2.49 0.14 0.015 6 0.56 1.34 0.56 0.70 2.04 2.59 0.14 0.015 8 0.56 1.34 0.59 0.73 2.07 2.66 0.15 0.015 Model R ṅ α × F × P × exp( ω × SC + λ ) 0.6895 RT α × F × P β × exp( ω × SC + λ ) 0.9523 RT α × F × SC × exp( × P + λ ) 0.9571 RT α × F × P × SC × exp( × P + ω × SC + λ ) 0.9485 RT α × F × P β × exp( × SC λ ) 0.6996 RT α × F × P β × exp( × SC λ ) 0.9550 RT α × F × P × exp( × SC λ + ) 0.6940 RT α × F × P β × exp( × SC λ + RT ) 0.9668
	4	0.07	0.015	0.43	0.055	0.03	0.014	2.98E-06	5.77E-07
	6	0.20	0.024	0.68	0.027	0.11	0.020	8.40E-06	5.77E-07
	8	0.29	0.032	0.86	0.020	0.19	0.026	1.24E-05	5.77E-07
	10	0.38	0.040	0.90	0.012	0.25	0.032	1.61E-05	5.77E-07
	12	0.46	0.048	0.93	0.006	0.31	0.037	1.95E-05	5.77E-07
					200 201 202 203 204 217

  Dans ce travail, la production in-situ d'hydrogène (pur) à partir de vapo-reformage d'éthanol (ESR) dans un réacteur catalytique à membrane (MR) a été étudiée. Un mélange d'éthanol pur et distillé a été utilisé comme combustible. Le réacteur est constitué d'un catalyseur Pd-Rh/CeO2 et d'une membrane Pd-Ag: l'ensemble est désigné par « reformeur ». Les expériences sur ce reformeur ont été effectuées dans diverses conditions de fonctionnement: température, pression, débit de combustible et rapport molaire de l'eau-éthanol (rapport SC). La performance du réacteur catalytique à membrane (CMR) a été étudiée en termes de facteur de production d'hydrogène théorique, d'efficacité de production de l'hydrogène et de la part d'hydrogène récupérée. L'évaluation thermodynamique du reformeur a été présentée. L'analyse exergétique a été réalisée sur la base des résultats expérimentaux visant non seulement à comprendre la performance thermodynamique du reformeur, mais aussi d'introduire l'application de l'analyse exergétique dans les études CMRs. L'analyse exergétique a fourni des informations importantes sur l'effet des conditions d'exploitation et les pertes thermodynamiques, et a donné lieu à la compréhension des meilleures conditions de fonctionnement. Outre les évaluations expérimentales et thermodynamiques du reformeur, la simulation de la dynamique de la production d'hydrogène (perméation) a été effectuée comme la dernière étape pour étudier l'applicabilité d'un tel système dans le cadre d'une utilisation finale réelle, qui peut être l'alimentation d'une pile à combustible. La simulation présentée dans ce travail est semblable aux ajustements de débit d'hydrogène nécessaires pour régler la charge électrique d'une pile à combustible répondant à des besoins variables.

ℎ = ℎ -ℎ -( -)4.9

Acknowledgments

I would like to express my greatest thanks to the following persons and organizations for their contribution:

The SELECT+ EMJD Doctoral Program for the scholarship and making it possible for me to pursue my academic career through a PhD.

performance of the membrane, differently at different operating conditions. The static model of the catalytic zone (based on the Arrhenius law) was used to fit the molar flow rate of the species present in the retentate gas, i.e. CO, CO 2 , CH 4 , H 2 , and H 2 O (to calculate the partial pressure of hydrogen at the retentate side).

Regarding eq. 5.6, the activation energy (E a ) was taken from the work by Papadias et al. [START_REF] Papadias | An analytical and experimental investigation of high-pressure catalytic steam reforming of ethanol in a hydrogen selective membrane reactor[END_REF] as they used the same membrane module as the one used in this work, with the same characteristics and synthetized by the same manufacturer (REB Research & Consulting [179]). Therefore, the term δ in eq. 5.6 was directly calculated, which is equal to 54.9

[m -1 ]. Then, the term 'Q 0 ' was obtained firstly from the experimental results (Q ), and then modeled (Q

) by means of a static model as a function of the reactor pressure (P) and fuel flow rate (F):

5.8

= × × ( × ) 5.9

Where ' ' is the fitting parameter. P H2,r in eq. 5.8 is obtained via eq. 5.7 by using the modeled values of the molar production rate of ESR products to calculate the hydrogen fraction in the catalytic zone. In fact, this factor (Q ) was used to fit the results of the Sieverts' law based model to the experimental ones.

The second model, named as "model 2", is a black box model, which was introduced as Model 2 in Table 5.3. This model is a function of the reactor pressure and the molar production rate of hydrogen (ṅ ) obtained by the steady state model developed for the catalytic zone (Model 5 presented in Table 5.2):

is the pure hydrogen production rate obtained via the black box model.

Accordingly, the hydrogen permeation rate at steady state conditions was modeled by means of model 1 and model 2, to be used in the simulation of the dynamics of hydrogen permeation rate at transient conditions, i.e. between two steady state points.

• Isothermal dynamic simulation of the permeation zone

Prior to the dynamic simulation of the permeation zone, the reactor pressure was modeled in the case of pressure set point adjustment during which the pressure valve of the reforming systems acts on the retentate gas flow rate to adjust to a higher or lower pressure.

The ideal gas law in the form of PV = W was used to model the pressure of the reactor.

Chapter 6

General conclusion

On-site pure hydrogen production in a catalytic membrane reactor by ethanol steam reforming Moreover, the advanced aspect of the control of the reforming system in connection with a real end user (for example a PEMFC) can be investigated and simulated.

163

Appendix A -Uncertainty Analysis

Reporting the measurement results is complete only when it is accompanied with uncertainty analysis. Because of inevitable errors during a measurement, a range of uncertainty is given to indicate a specific margin of doubt during which the measurement can be reliable. In fact, uncertainty of measurement states the doubt about the measurement results, and it differs from measurement error, which is the difference between the true value and the measured one.

In this work, uncertainty analysis is presented for experimental results and exergetic evaluation of the reforming system. Uncertainty of the measurement is rooted in the measuring instrument and the operator who is measuring a property. Accordingly, the uncertainty is calculated taking into account two sources of errors:

• Human error: the error in reading values during fuel blend preparation (water + ethanol) and measurement of several experimental properties and results through experiments in the laboratory • Error in laboratory equipment: beakers, graduated cylinder, flask, and so on, used in the preparation of fuel mixtures, together with the mass flow meters, the pressure valves, and specially the Gas Chromatograph.

The human error is considered in two stages. Firstly, the preparation of the fuel mixture where specific volume of water had to be mixed with a specific volume of ethanol to obtain a specific S/C ratio. It is assumed that a maximum human error of ± 0.5 ml is inevitable when measuring the value of the volume of the liquids by means of a graduated cylinder. Secondly, since the waste gas flow rate (retentate) is measured by mean of a bubble meter and a stopwatch, a human error of ±1 ml is considered in reading the volume of the gas, which has passed a certain volume of the bubble meter in a specific time.

The errors of laboratory measurement equipment are listed as:

• Selectivity of the gases measured by Gas Chromatograph (GC): ±5% absolute • Mass flow meter of the pure hydrogen stream: ±1.5 ml/min = ±10 -6 mole/s When the measured value by an instrument lies in a certain and known domain (±a), the uncertainty of the measurement is calculated as [START_REF] Bell | A Beginner's Guide to Uncertainty of Measurement[END_REF]219]:

Where u(c) is the standard uncertainty of the measured value. This is applicable in the case of the GC and mass flow meter of pure hydrogen.

The exergy rate of the retentate gas is composed of the exergy rate of each component (i), which exists in the stream, i.e. H 2 , CO, CO 2 , CH 4 , plus condensing water:

) A76

And finally:

A.2.9 Uncertainty of exergy efficiency:

Appendix B -general and specific models

The results of the static models given in chapter 4 are presented here in two sections, i.e. the general models (B.1) and specific models (B.2).

The fitting parameters of the models together with the coefficient of determination (R 2 ) of the species (H 2 , CO, CO 2 , CH 4 , and H 2 O) are given in the tables. The average coefficient of determination of the modeling ( ) is the arithmetic mean of the R 2 values of the five species:

The values of ω in the case of CO 2 and are positive, proving that at higher S/C ratio, more CO 2 is formed, and naturally, there is excess water fed into and leaving the reactor. The smallest value of ω belongs to CO that has the lowest selectivity among the ESR products. Very high values of can be ascribed to the significant effect of the temperature on the ESR reaction promotion and hydrogen permeation rate.

Considering θ, the negative values in the case of H 2 , CO, and CH 4 proves that the production rates of these species are conversely proportional to S/C ratio, which is in agreement with the experimental results. The high positive value in the case of H 2 O represents the effect of excess water at higher S/C ratios. The value of in the case of Hβ is positive and higher since the most effective factor on the hydrogen production (and permeation) is the operating pressure (and the partial pressure of hydrogen in the reactor). Very high values of can be ascribed to the significant effect of the temperature on the ESR reaction promotion and hydrogen permeation rate.

In this model, the combination of the effects of the pressure and S/C ratio can be seen. The values of ×P are very large (40000<P<1β0000 Pa) compared to the values of ω (1.6<SC<3), which is attributed to the dominant effect of pressure on the production rate of species. The values of place in a wild range similar to the values of α, which can be due to the combined effects of the operating conditions considered in this formulation. As seen, in the case of CH 4 and H 2 O, is very large and positive, while the values of α are some orders of magnitude smaller in comparison with the other three species.

The value of α is very large in the case of Hβ, which can be attributed to the effect of pressure (in the pre-exponential term) on the production rate of hydrogen. The values of SC are negligible ( <1) compared to , which can be interpreted as the minor effect of S/C ratio in comparison with temperature.

The value of is positive and larger in the case of hydrogen, which proves that the production rate of hydrogen in the CMR is directly proportional to the operating pressure. The values of SC are negligible ( <1) compared to , which can be interpreted as the minor effect of S/C ratio in comparison with temperature.

B.2 Specific Models

The specific models were developed at T=923 K.

The positive values of θ in the case of CO 2 and H 2 O can be attributed to the fact that the ESR reaction is promoted better at higher S/C ratio, probably due to the availability of excess water. The positive and larger value of in the case of hydrogen proves that the production rate of hydrogen in the CMR is directly proportional to the operating pressure.

Although the value of ω is higher in the case of CO, the value of is several orders of magnitude smaller than other species.

The positive values of θ in the case of CO 2 and H 2 O can be attributed to the fact that the ESR reaction is promoted better at higher S/C ratio, probably due to the availability of excess water. The positive and larger value of in the case of hydrogen proves that the production rate of hydrogen in the CMR is directly proportional to the operating pressure. the same explanation can be given in the case of ω. Considering the value of α × F β , the lower value is obtained in the case of H 2 , which can be attributed to the significant effect of pressure, which appears in the exponential term.

Mots clés

Réacteur à membrane, vapo-reformage de l'éthanol, teneur en hydrogène, récupération d'hydrogène, analyse exergétique, rendement exergétique, rendement thermique, modèle statique, simulation dynamique, loi de Sieverts

Abstract

In this work, in-situ production of fuel cell grade hydrogen (pure hydrogen) via catalytic ethanol steam reforming (ESR) in a membrane reactor (MR) was investigated. A mixture of pure ethanol and distilled was used as the fuel. ESR experiments were carried out over a Pd-Rh/CeO2 catalyst in a Pd-Ag membrane reactornamed as the fuel reformerat variety of operating conditions regarding the operating temperature, pressure, fuel flow rate, and the molar ratio of water-ethanol (S/C ratio). The performance of the catalytic membrane reactor (CMR) was studied in terms of pure hydrogen production, hydrogen yield, and hydrogen recovery. Thermodynamic evaluation of the CMR was presented as a supplement to the comprehensive investigation of the overall performance of the mentioned pure hydrogen generating system. Exergy analysis was performed based on the experimental results aiming not only to understand the thermodynamic performance of the fuel reformer, but also to introduce the application of the exergy analysis in CMRs studies. Exergy analysis provided important information on the effect of operating conditions and thermodynamic losses, resulting in understanding of the best operating conditions. In addition to the experimental and thermodynamic evaluation of the reforming system, the simulation of the dynamics of hydrogen production (permeation) was performed as the last step to study the applicability of such a system in connection with a real end user, which can be a fuel cell. The simulation presented in this work is similar to the hydrogen flow rate adjustments needed to set the electrical load of a fuel cell, if fed online by the studied pure hydrogen generating system.