Professeur Catherine Azzaro-Pantel

Professeur Marco Baratieri

Professeur Ange Nzihou

Chlorinated contaminants mitigation dur ing pyro-gasification of wastes using CaO reactant: Experimental and life cycle assessment et discipline ou spécialité

Keywords: Municipal solid waste, Pyro-gasification, HCl mitigation efficiency, CaO reactant, In-situ dehydrochlorination, Life cycle assessment

Chapter 1

To complete the aforementioned works, life cycle assessment (LCA) of three typical pyro-gasification and incineration processes is conducted to compare their overall environmental sustainability. Moreover, pyro-gasification-based WtE systems with different dehydrochlorination strategies are further modeled: 1) conventional gasification system; 2) novel gasification coupled with ex-situ high temperature dehydrochlorination system; and 3) hypothetical gasification coupled with in-situ dehydrochlorination system. The obtained results could be applied to optimize the current waste pyro-gasification systems, with special focus on developing strategies for in-situ dehydrochlorination purpose.

Background and literature review

The objective of this chapter is to provide insights and background on current research progress: HCl release from MSW thermal conversion and its control methods; role of CaO reactant on in-furnace tar cracking and upgrading; and LCA methodology and its application on MSW thermal conversion systems. With the overall aim, the structure of this chapter is set as follows:

-Section 1.1 presents the background and explains the motivation of the present study; -Section 1.2 introduces the presence of chlorine in MSW and HCl release from MSW thermal conversion process; besides, the state-of-the-art regarding dehydrochlorination methods are summarized; -Section 1.3 focuses on the introduction of reactivity of CaO on pyro-gasification syngas quality enhancement and tar cracking and upgrading; -Section 1. 4 presents the principle and framework of LCA as evaluation tool for waste system, and its current status application in WtE technologies; -Section 1.5 details the structure and contents of the thesis.

Background and motivation of the thesis

It has to be acknowledged that with the booming of social economy and urbanization all around the world, municipal solid waste (MSW) generation increased dramatically in the past several decades, especially in developing countries. According to the world bank [1], approximately 1.3 billion tons of MSW is generated annually in 2016; and this value is expected to reach 2.2 billion tons by 2025 according to their estimations. Figure 1. 1 presents the MSW generation amount in China and France during 2007-2016 [1]. As illustrated, MSW generations in France is quite stable in this period, ranging from 34.1 million tons to 34.8 million tons; however, MSW generation in China increased rapidly from 152.2 million tons in 2007 to 203.6 million tons in 2016, with a quite high annual increasing of ca. 3.3%. Although the total MSW generation amount in China is approximately 5 times higher than that in France, MSW generation per capita in China is nearly 1/5 of that in France. This reflects the fact that developed countries tend to generate more MSW per person as compared to developing Countries. However, due to the huge population in China, the total generation amount of MSW is very high and keeps increasing. [1] Under such circumstance, the treatment of MSW has become one of the biggest challenges facing with by the modern societies, owning to its huge amount and serious consequences if not properly treated (for example, air emissions, soil contaminations, groundwater pollution, and even serious public health risks). In recent years, the sanitary landfill is the most popular technology used for MSW treatment in many Countries, because it is the least expensive alternative [2,3]. However, concerns with respect to MSW have increasingly been shifted from disposal to integrate waste management. The waste hierarchy, as defined by European Directive 2006/12/EC [4] and updated by Directive 2008/98/EC [5] in 2008, ranks the waste management strategies from the top priority of waste prevention, followed by waste reuse, recycling, recovery, and the final option waste disposal. The hierarchy is mainly based on the principle that which waste treatment option is best for the environment. Figure 1. 2 illustrates the preference and contents of the waste hierarchy according to the Department for Environment Food & Rural Affairs of the U.K. [6].

Figure 1. 2 Preference and contents of the waste hierarchy as regulated by European

Directive, data derived from [6] In regard to the waste treatment strategies adopted by different regions, Figure 1. 3 summarizes the share of treatment methods in the E.U. Countries during 2004 to 2015 [7].

As the figure indicates, the most popular MSW treatment methods include landfill, incineration and recycling. And in European countries, recycling has performed an important role in waste management systems. Besides, incineration has also played an increasingly important role in waste management over the past decades, mainly due to the advantages of significant waste reduction, complete disinfection, as well as energy recovery [8]. By incineration, approximately 70% of weight and 90% of volume of MSW are reduced and around 20% of the energy content is able to be recovered in the form of electricity and/or heat [9][10][11]. Data from the literature revealed that in 2012, approximately 14.5 million tons of MSW (more than 28% of total generated MSW in France) were incinerated in 126 MSW incineration plants [11].Therefore, it is well recognized as an efficient energy recovery method as depicted also in the E.U. waste hierarchy. [7] Although MSW incineration becomes more and more popular in recent years, problems are still existing for its popularizing in some Countries and regions, mainly concerning the harmful emissions, especially heavy metals, dioxins and furans, and greenhouse gas emissions [12]. Moreover, due to the relatively high content of chlorine (Cl) in MSW, hydrogen chloride (HCl) is usually generated as a consequence of high temperature thermal conversion. The presence of HCl in incineration flue gas may cause serious problems like high temperature corrosion, toxic organic contaminants (mainly dioxins, furans, and VOCs) formation, acidification, etc. Besides, due to the corrosion potential aroused by HCl gas, the temperature and pressure of the steam for electricity generation are strictly limited at 400 o C and 40 bar, resulting in the relatively low electricity generation efficiency of MSW incineration plants. According to the statistical results by Beylot et al. [11], the net electricity generation efficiency is as low as 14.96% when only electricity is recovered from French MSW incineration plants. Therefore, chlorine-content contaminants in incineration flue gas must be removed in order to lower the environmental burdens as well as to improve the electricity generation efficiency. Currently, HCl is usually removed together with acid gases by air pollution control (APC) systems, which basically works in temperature range of lower than 300 o C. Although the removal efficiency of HCl by APC systems is high enough, HCl has already served as corrosive agent and chlorine source for polychlorinated dibenzo-p-dioxins and polychlorinated dibenzofurans (PCDD/Fs) formation during flue gas cooling down under such conditions. Therefore, HCl removal at higher temperature (i.e. before it causes problems such as corrosion, Cl-source for PCDD/Fs formation, etc.) is of great importance.

Figure 1. 4 Conceptual diagram of Kymijärvi II RDF gasification power plant, located in Lahti

region, Finland, data acquired from [13] MSW pyro-gasification technology offers the possibilities to achieve the aforementioned target. MSW pyro-gasification technology is defined as thermochemical conversion treatment of MSW to intermediate products (such as produced gas, tar, and char) in the absence of oxygen (pyrolysis) or in partial combustion conditions (gasification). Compared to si gle-step i i e atio , p o-gasification-ased WtE te h i ue, i.e. t o-step o idatio , makes it possible to decrease dioxins and NO x emission as a result of the reducing atmosphere [14,15]; and the intermediate products could be cleaned to remove contaminants prior to the subsequent combustion thanks to the two separated conversion steps. The idea has been put into practice by Kymijärvi II power plant, located in Lahti region, Finland [16]. In this novel gasification plant, the solid recovered fuel (SRF) is firstly gasified in a circulating fluidized bed (CFB) gasifier at temperature of 850-900 o C; then the produced combustible gas is cooled down to around 400 o C and then undergoes hot gas cleanup devices to remove corrosive chlorine-content contaminants and particulate matters (PM) by ceramic candle filters; and finally, the purified syngas is combusted to generate a steam with high parameters (temperature of 540 o C and pressure of 120 bar) in the boiler and thus gives a higher overall power production efficiency [13,17]. Figure 14 exhibits the conceptual diagram of Kymijärvi II power plant [13].

As discussed, the improvement of both energy efficiency enhancement and environmental impacts reduction is quite obvious by the advanced Kymijärvi II MSW gasification power plant thanks to the hot syngas purification stage. However, energy degradations still exist when cooling down the produced syngas before being purified; besides, the ex-situ removal of chlorine-content contaminants makes the whole gasification-based WtE system more complex and increases the operation difficulties.

Realizing these facts, the in-furnace dechlorination by using CaO reactant is proposed and investigated in the present work. Accordingly, the motivation of the present thesis is to investigate the in-situ removal of chlorine-content contaminants by adding reactants during MSW pyro-gasification process.

The primary idea of the present work, as illustrated in Figure 1.5, is to prevent the chlorine-content contaminants from discharging into the produced gas by using in-situ CaO reactant. Therefore, effect of CaO for the removal of the released HCl is investigated and discussed in this dissertation. Moreover, effect of CaO reactant on pyrolytic tar upgrading is also studied.

HCl release and control methods during waste thermal conversion

It is well recognized that chlorine (Cl) content in MSW is relatively high and ranges from 800 to 2500 mg/kg according to Jhong-Lin Wu et al. [18]. The presence of chlorine in MSW is mainly in the form of inorganic chlorine such as NaCl or KCl and organic chlorine such as PVC or rubber [19]. However, no matter inorganic chlorine or organic chlorine, the main Cl-content contaminants produced from MSW thermal treatment is HCl. According to the literature, HCl concentration in the incinerator out flow can be as high as 800-2000 mg/Nm 3 .

Besides, enrichment of Cl in incineration fly ash due to condensation is serious and its concentration is typically from 1.8 to 9.1 wt. % [20]. As aforementioned in previous section, the presence of HCl in incineration flue gas causes significant adverse effects on both electricity generation efficiency and environment burdens. Therefore, thorough investigation of HCl release characteristics as well as its control methods is of great importance.

Presence and release of chlorine from MSW 1.2.1.1 Chlorine contained in inorganic source

As one of the major presence form of chlorine in MSW, inorganic chlorine plays an important role in MSW management. The presence of inorganic chlorine in MSW is mainly related to alkali chlorides (for example NaCl in food waste and KCl in agricultural residues) [21][22][23][24][25]. Among MSW fractions, food waste is the most important inorganic chlorine sources, according to Ma et al. [23], approximately more than 45% of the total chlorine contained in MSW comes from food waste. Although inorganic chlorine takes a high proportion of the total chlorine in MSW, it is believed that only a small amount of chlorine is converted directly from alkali chlorides to HCl or other Cl-content contaminants [26]. The reason may be attributed to the relative high binding energy of alkali chlorides (for example, NaCl 787 kJ/mol, and KCl 717 kJ/mol) [22,27]. Takasuga et al. reported that NaCl does not produce HCl itself in the case of direct heating [28]. However, the presence of water, sulfur, O 2 , and even some inorganic materials like SiO 2 and Al 2 O 3 facilitates the formation of HCl from alkali chlorides [29,30]. The formation routes can be described as follows (taking NaCl as example): (1.7)

2 NaCl H O NaOH HCl    (1.1)

Chlorine contained in organic source

Apart from inorganic chlorine content, organic chlorine is also important or even more important in regard to chlorine-content contaminants formation during waste thermal conversion. Generally, organic chlorine contained in MSW are mainly in the forms of chlorinated plastics, for example, polyvinyl chloride (PVC), chlorinated polyethylene (CPE), polyvinyl dichloride (PVDC), electronic waste, textiles, leathers, rubbers [START_REF] Coda | Behavior of chlorine and enrichment of risky elements in bubbling fluidized bed combustion of biomass and waste assisted by additives[END_REF][32][33], among which, PVC is the most representative organic chlorine source in MSW and have attracted worldwide concerns in the last decades [34,35]. Theoretically, chlorine content in PVC is as high as 56 wt. % of the total mass. Unlike inorganic chlorine, organic chlorine in PVC releases [ [36][START_REF] Slapak | Determination of the pyrolytic degradation kinetics of virgin-PVC and PVC-waste by analytical and computational methods[END_REF][38][39]. Therefore, organic chlorine contained in MSW is easy to be released and causes serious problems.

Possible hazards from HCl emission during MSW thermal conversion 1.2.2.1 Cause of acid rain in local area

Acid rain is defined as the accumulation of atmospheric acidic substances in wet or dry conditions in the ecosystem with pH lower than 5.6 [40,[START_REF] Liang | Effects of simulated acid rain on soil respiration and its components in a subtropical mixed conifer and broadleaf forest in southern China[END_REF]. Acid rain is one of the most serious environmental issues facing by the world. Especially in China, approximately 40% of the main land is affected by acid rain, and China is the third largest acid rain area, ranking only after Europe and North America [START_REF] Liang | Effects of simulated acid rain on soil respiration and its components in a subtropical mixed conifer and broadleaf forest in southern China[END_REF]. 4) may cause harmful effects on soil, forests and water, data obtained from [START_REF] Agency | What is Acid Rain?[END_REF] HCl is considered the third largest source of acid rain formation (the first and second sources are SO 2 and NO x , respectively). The formation path of acid gas from HCl is depicted in Figure 1.6 [START_REF] Agency | What is Acid Rain?[END_REF] Different from SO 2 and NO x , which may be transported to a quite long distance before falling to the ground, HCl is highly soluble in rainwater; therefore, it is a local pollution source and will affect directly the local surroundings of the emission source [START_REF] Williams-Jones | Hazards of volcanic gases. The Encyclopedia of Volcanoes[END_REF].

Typically, the potential harm of acid rain is summarized as follows [START_REF]Effects of Acid Rain[END_REF]:

-Firstly, acid rain significantly affects the aquatic ecosystems. The low pH value of acid rain causes acidification of the water, leading to structural and functional imbalances of the aquatic ecosystem. On the one hand, the growth and reproduction of fish will be affected; on the other hand, it leads to the change of composition and acid resistance of water, for example, algae and fungi content may increase, and other animals and plants may decrease simultaneously.

-Secondly, acid rain affects the terrestrial ecosystems. The impacts are mainly on two aspects: soil and vegetation. In the case of soil, acid rain may change the physical and chemical properties of soil. And also it may activate some harmful and/or toxic elements contained in soil. In terms of vegetation, the fell of acid rain directly destroy the vegetation. Besides, the impact on soil will also indirectly affect the growth of plants.

-Thirdly, acid rain causes damages on buildings and artifacts. Acid rain will react with metals, concrete and other materials under certain conditions, thus affect buildings and artifacts exposed to rainwater.

-Last but not least, the formation of acid rain will have an adverse impact on human health. On the one hand, it directly irritates our skin and may cause various respiratory diseases such as asthma; on the other hand, it will result in the inrush of harmful elements into the water, causing water pollution, thereby pollutes food and fish that indirectly affect human health.

High temperature corrosion of the boiler

As aforementioned, the presence of chlorine in MSW and the formation of HCl significantly restrict the steam quality for electricity generation. Typically in waste incinerator, the temperature and pressure of the produced steam are 400 o C and 4.0 MPa, which is much lower than that of coal-fired power plants at around 550 o C, 22 MPa for traditional subcritical power plants, or 565 o C, 24.3 MPa for supercritical power plants, resulting in a quite low electricity generation efficiency of around 20% [11,[START_REF] Li | Clean coal conversion processes-progress and challenges[END_REF]. However, this value for coal-fired power plant is in the range of 33-37% and 37-40% for subcritical and supercritical power plants, respectively [START_REF] Li | Clean coal conversion processes-progress and challenges[END_REF]. High temperature corrosion of the equipment by HCl is responsible for this problem. The formed HCl gas may cause serious damages in incineration chamber, water wall tube and super heater through chemical reactions with metal/metal oxide [START_REF] Haleem | Factors affecting the corrosion behaviour of aluminium in acid solutions. I. Nitrogen and/or sulphur-containing organic compounds as corrosion inhibitors for Al in HCl solutions[END_REF][START_REF] Wang | Affection of High Sulfur and Naphthenic Acid Crude Oil Acted on Refinery Equipment[END_REF][START_REF] James | Effect of coal chlorine on the fireside corrosion of boiler furnace wall and superheater/reheater tubing[END_REF].

The principle of high temperature corrosion occurred on the surface of the equipment are illustrated in Figure 1.7 [START_REF] Ting | Study of calcium oxide dechlorination in biomass fluidized bed gasifier[END_REF].

1.15

Accordingly, the high temperature corrosion process related to HCl or Cl 2 are described as: 1) direct chemical reactions between HCl/Cl 2 with metal/metal oxides to form metal chlorides which melts at low temperature, resulting in the loss of materials; 2) the formation of low-melting eutectic mixtures by metal/metal oxides and low-melting chlorides such as NaCl and sulfate, thereby accelerating the high-temperature corrosion process; 3) diffusion of chlorine into the defects, such as grain boundaries, voids, and etc., of the metal, causing internal chlorination of the metal matrix and accelerating the corrosion rate. Therefore, once HCl or Cl 2 exists in the flue gas, the high temperature corrosion reactions will keep going.

Possible chemical reactions are summarized in Table 1.1 [START_REF] James | Effect of coal chlorine on the fireside corrosion of boiler furnace wall and superheater/reheater tubing[END_REF][START_REF] Ting | Study of calcium oxide dechlorination in biomass fluidized bed gasifier[END_REF].

Chlorine source for formation of polychlorinated dibenzo-p-dioxins and polychlorinated dibenzofurans (PCDD/Fs)

Previous studies revealed that the formation of PCDD/Fs (i.e. dioxins) during MSW thermal conversion is mainly through homogeneous and heterogeneous pathways [20,[START_REF] Altarawneh | Mechanisms for formation, chlorination, dechlorination and destruction of polychlorinated dibenzo-p-dioxins and dibenzofurans (PCDD/Fs)[END_REF][START_REF] Vermeulen | Indication of PCDD/F formation through precursor condensation in a full-scale hazardous waste incinerator[END_REF][START_REF] Rathna | Recent developments and prospects of dioxins and furans remediation[END_REF][START_REF] Collina | The rate-determining step in a low temperature PCDD/F formation from oxidative breakdown of native carbon in MSWI fly ash[END_REF]. The homogeneous pathway occurs at temperature between 400 and 800 °C, involving the conversion from structurally related precursors (such as chlorinated phenols and chlorobenzenes) to dioxins [START_REF] Altarawneh | Mechanisms for formation, chlorination, dechlorination and destruction of polychlorinated dibenzo-p-dioxins and dibenzofurans (PCDD/Fs)[END_REF][START_REF] Rathna | Recent developments and prospects of dioxins and furans remediation[END_REF]. In regard to the heterogeneous pathway, there are two formation mechanisms [20,[START_REF] Zhang | De novo synthesis of dioxins: a review[END_REF]: one is called de novo process, which occurs at temperature range of 200 -400 °C and proceeds through catalytic chlorination and oxidation reactions of carbonaceous materials from elemental C, H, O, and Cl directly to dioxins; and the other one is named precursor routes, which also takes place at the temperature range of 200 -400 °C and refers to the formation of dioxins from structurally related species, mainly chlorinated organic compounds like chlorobenzene, chlorophenol, polychlorinated biphenyls, and so on.

As indicated, no matter homogeneous and heterogeneous dioxins formation pathways, chlorine and chlorinated-contaminants play important roles. With special concerns to the heterogeneous pathways, as aforementioned in general introduction section, the temperature window for heterogeneous synthesis overlaps with the working temperature of APC systems, which means before being removed from incineration flue gas, the heterogeneous synthesis of dioxins is occurred. Realizing this fact, high temperature removal of chlorine-content contaminants is very meaningful for the inhibition of dioxins formation during MSW thermal treatment. Due to the potential harmfulness of HCl on both human health and the environment, HCl emission is strictly regulated worldwide. HCl emission standard in China [START_REF]Pollution Control Standard for Municipal Solid Waste Incineration (GB 18485-2014[END_REF], the European Union [START_REF] Directive | Directive on the incineration of waste[END_REF], and the United States [START_REF]Commercial and Industrial Solid Waste Incineration Units: Reconsideration and Final Amendments; NonHazardous Secondary Materials That Are Solid Waste[END_REF][START_REF]Emission Guidelines for Existing Small Municipal Waste Combustion Units[END_REF] are presented in Table 1.2.

Current HCl emission regulations

HCl control methods

In order to meet the strict requirements of the emission regulations as well as to reduce air emission caused by waste thermal treatment, several control methods are proposed and applied to reduce HCl emission. The control method can be summarized as: 1) HCl emission control before waste thermal treatment by MSW pre-treatment; 2) HCl emission control during waste thermal treatment by in-furnace reactant; and 3) HCl emission control after waste thermal treatments by air pollution control facilities. The general principles of HCl elimination method for waste thermal conversion are presented as follows.

HCl control before waste thermal treatment

The most effective method to prevent the formation of chlorine content contaminants from MSW thermal treatments is to remove the chlorine contained in raw MSW. Therefore, pre-treatment of waste before thermal conversion is an effective HCl emission control strategies. Pre-treatment methods such as water washing [START_REF] Deng | Effect of water washing on fuel properties, pyrolysis and combustion characteristics, and ash fusibility of biomass[END_REF][START_REF] Jensen | Removal of K and Cl by leaching of straw char[END_REF][START_REF] Bakker | Feasibility of collecting naturally leached rice straw for thermal conversion[END_REF], source-separated collection [START_REF] Nasrullah | Elemental balance of SRF production process: solid recovered fuel produced from commercial and industrial waste[END_REF][START_REF] Sarc | Production, quality and quality assurance of Refuse Derived Fuels (RDFs)[END_REF][START_REF] Velis | Solid recovered fuel: materials flow analysis and fuel property development during the mechanical processing of biodried waste[END_REF][START_REF] Garcés | Evaluation of the potential of different high calorific waste fractions for the preparation of solid recovered fuels[END_REF], and low-temperature thermal treatment is commonly used ways to lower the chlorine content:

-Water washing Pre-treatment by water washing is effective to remove most of the water-soluble choline contained in MSW. As previously mentioned in Section 1.1.1, inorganic chlorine content in MSW is mainly in the form of NaCl and KCl; and it is water-soluble. Deng et al. found that by washing with deionized water, more than 90% of chlorine content in rice husk can be removed [START_REF] Deng | Effect of water washing on fuel properties, pyrolysis and combustion characteristics, and ash fusibility of biomass[END_REF].

-Sorting Water washing treats mainly inorganic chlorine content in MSW. For organic chlorine, the most effective pre-treatment method is source-separated collection and waste recycling, especially for chlorine-content plastics. Unfortunately, PVC, the most important organic chlorine source in MSW, is not currently recycled in many Countries and regions and mixed in the non-recyclable plastics according to Kikuchi et al. [START_REF] Kikuchi | Grouping of mixed waste plastics according to chlorine content[END_REF] and Themelis at al. [START_REF] Themelis | Chlorine Sources, Sinks, and Impacts in WTE Power Plants[END_REF], resulting in the high organic Cl content in MSW.

-Low-temperature thermal treatment It is reported that chlorine content in organic source, such as PVC, releases at low temperature range of 200-360 o C [START_REF] Kim | Pyrolysis kinetics of waste PVC pipe[END_REF][START_REF] Zhu | TG-FTIR analysis of PVC thermal degradation and HCl removal[END_REF]. Therefore, the idea of low-temperature thermal pre-treatment is to pre-pyrolyze the Cl content MSW before being thermally treated. And it is usually divided into two steps: (1) low temperature pyrolysis of raw MSW at 200-360 °C to release most Cl by dehydrochlorination reaction, and (2) thermal treatment of de-HCl MSW at higher temperatures.

HCl control after waste thermal treatment

Currently, the most important HCl control method is the use of APC systems after thermal treatment process to meet the emission standards [START_REF] Hammoud | Solid waste to energy strategy in Lebanon: Potential, technology and design[END_REF]. Therefore, HCl removal is usually accomplished together with water-soluble acid gases such as hydrofluoric acid (HF), sulphur dioxide (SO 2 ), and etc. The acid gas removal efficiency depends highly on the selected method and equipment, the flue gas composition, and the sorbents. Typically, the most commonly used technologies for acid gases removal are: (1) dry treatment method with calcium bicarbonate (NaHCO 3 ) or calcium hydroxide (Ca(OH) 2 ) injection; (2) wet scrubbing method; and (3) semi-dry treatment method with the use of Ca(OH) 2 slurry [START_REF] Ji | Municipal solid waste incineration in China and the issue of acidification: A review[END_REF][START_REF] Tian | Atmospheric pollution problems and control proposals associated with solid waste management in China: a review[END_REF][START_REF] Zhou | Development of an intelligent decision support system for air pollution control at coal-fired power plants[END_REF].

-Dry treatment method

Regarding to dry treatment method of acid gases in the flue gas, fine sorbents powder, mostly hydrated lime or NaHCO 3 , is directly injected into the furnace or the flue gas pipeline to remove the acid gases content. Among the sorbents, NaHCO 3 is a promising alternative sorbent, which has higher reaction rates, lower consumption, and produces fewer residues; however it has a relatively higher unit cost. Dry treatment method for acid gases with NaHCO 3 is based on the following reaction that take place at least 130 o C. The decomposition of NaHCO 3 produces Na 2 CO 3, which has a e high spe ifi su fa e area and can effectively react with HCl to form NaCl, thus immobilizing the generated HCl in the flue gas.

The sorbent is usually introduced through a nozzle into the transfer line or into a de-acidification reactor to realize adequate contact time of gas and sorbents. De-acidification reactions occur on the surface of the sprayed dry Ca(OH) 2 or Na 2 CO 3 particle to form solid chlorides. The produced solids, as well as unreacted sorbent, are then separated in the dust collector.

To meet the strict requirements of acid gases emission regulations, sufficient contact time of de-acidification sorbent and flues gas is essential. The reaction rates depend on the reaction temperature, the moisture content, the concentration of acid gases in the flue gas, and the physical properties (for example: particle size, pore distribution, specific surface area and etc.) of the sorbents.

-Wet treatment method

Generally, the wet scrubbing method has the best performance regarding to the term of acid gases removal efficiency. Meanwhile, the amount of the required reactants can be less when compared with dry treatment methods. Typically, wet scrubbing method can be divided into two stages: the first stage is to quench the flue gas to the temperature of 60-70 o C by spraying water, since most of the flue gas entering the wet scrubber is unsaturated; the second stage is the injection of the solution and/or slurry for the absorption of acid gases.

Wet scrubbers are usually placed after the particle matter removal equipment, to prevent their clogging by sludge.

Several low-soluble calcium salts are generated during the absorption reaction of acid gases (for example chlorides, sulphates, and phosphates...). Thus, the circulating water has to be purified, concentrated and filtered, in order to avoid salt depositing. Meanwhile, the produced wastewater needs to be treated before being discharged.

-Semi-dry treatment methods

Semi-dry scrubber system is the most preferable treatment methods applied in MSW incineration plants in China. Generally, by using the semi-dry acid gases treatment methods, the sorbent slurry is sprayed into the reaction tower by nozzles or rotary atomiser to produce fine droplets. The evaporation of the moisture content slightly increases the humidity and decreases the temperature of the flue gas, resulting in an improvement of acid gases removal efficiency. To sum up, the advantages and disadvantages of the aforementioned three HCl removal methods are summarized in Table 1. 4. Wastewater generation, containing chlorides and heavy metals;

Off-gas must be reheated to above 140 °C to prevent the formation of a white plume at the stack outlet, which is quite energy consuming.

Investments and operating costs are quite high.

Semi-dry treatment methods

No wastewater produced;

Operating temperature is usually 140 °C or higher(thus spray de-superheating system is not essential);

Water consumption is lower than wet scrubbers and the exhaust gas can be discharged directly after removal of PM.

Higher operation costs because of reagents used;

Higher maintenance costs for dust removal equipment

HCl control during waste thermal treatment

The use of in-furnace reactant is the main strategy for the removal of released HCl during waste thermal treatment. By using the sorbents, Cl is captured by either physical adsorption or chemical absorption and is immobilized into solid residues. Different kinds of sorbent have been examined for the in-furnace removal of released HCl, for example alkaline and alkaline earth metal content sorbents, CaO [START_REF] Zhu | TG-FTIR analysis of PVC thermal degradation and HCl removal[END_REF][START_REF] Sun | Experimental study of HCl capture using CaO reactants: activation, deactivation, reactivation, and ionic transfer mechanism[END_REF][START_REF] Dou | High-temperature HCl removal with sorbents in a fixed-bed reactor[END_REF][START_REF] Bie | Reaction mechanism of CaO with HCl in incineration of wastewater in fluidized bed[END_REF][START_REF] Cho | Air gasification of PVC (polyvinyl chloride)-containing plastic waste in a two-stage gasifier using Ca-based additives and Ni-loaded activated carbon for the production of clean and hydrogen-rich producer gas[END_REF], Magnesium Oxide [START_REF] Dou | High-temperature HCl removal with sorbents in a fixed-bed reactor[END_REF][START_REF] Dou | A study on the reaction kinetics of HCl removal from high-temperature coal gas[END_REF],

CaCO 3 [START_REF] Zhu | TG-FTIR analysis of PVC thermal degradation and HCl removal[END_REF][START_REF] Dou | High-temperature HCl removal with sorbents in a fixed-bed reactor[END_REF][START_REF] Fontana | Municipal waste pyrolysis (2) chlorine capture by addition of calcium and sodium-based sorbents. Erdoel, Erdgas[END_REF][START_REF] López | Dechlorination of fuels in pyrolysis of PVC containing plastic wastes[END_REF], Ca(OH) 2 [START_REF] Zhu | TG-FTIR analysis of PVC thermal degradation and HCl removal[END_REF][START_REF] Fontana | Municipal waste pyrolysis (2) chlorine capture by addition of calcium and sodium-based sorbents. Erdoel, Erdgas[END_REF], NaHCO 3 [START_REF] Fontana | Municipal waste pyrolysis (2) chlorine capture by addition of calcium and sodium-based sorbents. Erdoel, Erdgas[END_REF][START_REF] Kong | Dry sorbent injection of sodium sorbents for SO2, HCl and mercury mitigation[END_REF]; and natural minerals, dolomite [START_REF] Cho | Air gasification of PVC (polyvinyl chloride)-containing plastic waste in a two-stage gasifier using Ca-based additives and Ni-loaded activated carbon for the production of clean and hydrogen-rich producer gas[END_REF][START_REF] Cho | Two-stage air gasification of mixed plastic waste: Olivine as the bed material and effects of various additives and a nickel-plated distributor on the tar removal[END_REF]84], olivine [START_REF] Cho | Air gasification of PVC (polyvinyl chloride)-containing plastic waste in a two-stage gasifier using Ca-based additives and Ni-loaded activated carbon for the production of clean and hydrogen-rich producer gas[END_REF][START_REF] Cho | Two-stage air gasification of mixed plastic waste: Olivine as the bed material and effects of various additives and a nickel-plated distributor on the tar removal[END_REF].

Among the sorbents, calcium-based reactants are the most popular choices due to abundant sources, relatively cheap and high HCl removal efficiency by reactions presented in Equation 1.17 -1.22 [START_REF] Bausach | Reaction of Ca (OH) 2 with HCl in the presence of water vapour at low temperatures[END_REF][START_REF] Xie | HCl removal using cycled carbide slag from calcium looping cycles[END_REF]. Weinell et al. [START_REF] Weinell | Hydrogen chloride reaction with lime and limestone: kinetics and sorption capacity[END_REF] studied the reaction between HCl and limestone at the temperature range of 60-1000 o C and concluded that the reaction is mainly controlled by product-layer diffusion according to unreacted shrinking core model; they also point out that at temperature higher than 500 o C, the absorption of HCl is limited by chemical reaction equilibrium. Gullett et al. modeled the reaction kinetics between Ca-based sorbents and HCl, and stated that at temperature range of 150-350 o C, the apparent activation energy is 28.1 kJmol -1 and it can be described as first order reaction with respect to HCl [START_REF] Gullett | Reaction kinetics of calcium-based sorbents with hydrogen chloride[END_REF].

CaO HCl CaClOH   (1.17) 2 2 2 CaO HCl CaCl H O    (1.18) 3 2 CaCO HCl CaClOH CO    (1.19) 3 2 2 2 2 CaCO HCl CaCl H O CO     (1.20)   2 2 Ca OH HCl CaClOH H O    (1.21)  
Accordingly, CaO is selected as the dehydrochlorination sorbent in this study. The gas-solid reaction mechanism is investigated by performing HCl breakthrough experiments and effect of in-reactor CaO on released HCl mitigation during MSW pyrolysis is also identified.

Effect of CaO on pyro-gasification syngas quality enhancement and tar cracking and upgrading

Besides serving as dehydrochlorination sorbents, the addition of CaO into MSW pyro-gasification reactor can also acts as a reactant for syngas quality enhancement and tar cracking [START_REF] Han | Influence of CaO additives on wheat-straw pyrolysis as determined by TG-FTIR analysis[END_REF][START_REF] Han | Hydrogen production via CaO sorption enhanced anaerobic gasification of sawdust in a bubbling fluidized bed[END_REF][START_REF] Zhou | Effect of calcium oxide on high-temperature steam gasification of municipal solid waste[END_REF][START_REF] Jordan | Effect of CaO on tar production and dew point depression during gasification of fuel cane bagasse in a novel downdraft gasifier[END_REF]. This section is dedicated to briefly introduce the recent research process on the reactivity of CaO on these aspects.

CaO as pyro-gasification syngas quality enhancement reactant

The addition of in-furnace CaO reactant takes effect via several pathways on syngas quality enhancement [START_REF] Dong | MSWs gasification with emphasis on energy, environment and life cycle assessment[END_REF]. On the one hand, CaO can be served as a good CO 2 acceptor and reacts with CO 2 to form CaCO 3 (CaO carbonization reaction as presented in Eq. 1.23). The reduction of CO 2 in the reaction system will lead to the enhancement of water gas shift reaction (as shown in Eq. 1.24) towards the direction to hydrogen production [START_REF] Florin | Enhanced hydrogen production from biomass with in situ carbon dioxide capture using calcium oxide sorbents[END_REF][START_REF] Florin | Hydrogen production from biomass coupled with carbon dioxide capture: the implications of thermodynamic equilibrium[END_REF][START_REF] Xu | Distinctive effects of CaO additive on atmospheric gasification of biomass at different temperatures[END_REF].

Dong et al. investigated the effect of in-furnace addition of CaO reactant during bio-char steam gasification, and concluded that CaO addition as in-situ reactant at 700 o C produces as high H 2 yield as that gasified at 800 o C without reactant [START_REF] Dong | Hydrogen-Rich Gas Production from Steam Gasification of Bio-char in the Presence of CaO[END_REF].

2 3 CaO CO CaCO   (1.23) 2 2 2 CO H O CO H    (1.24)
On the other hand, CaO can significantly reduce the organic impurities content (especially tar compounds) in the produced gas and converts them to combustible gases.

Both the removal of impurities and the enhancement of combustible gases generation are favorable for the downstream use of the produced gas. Eq. 1.25 conceptually introduces the impact of CaO on tar cracking.

         
(1.25)

CaO as pyrolysis tar upgrading reactant

Apart from being treated as impurities, tar generated from MSW pyro-gasification contains a large amount of high value added compounds, like benzene, toluene, xylene, and naphthalene (BTXN); therefore, the recovery of these chemicals from pyro-gasification tar is also one of the research hotspots [START_REF] Zhang | Upgrading of bio-oil from biomass fast pyrolysis in China: A review[END_REF][START_REF] Yan | Effects of low molecular compounds in coal on the catalytic upgrading of gaseous tar[END_REF][START_REF] Chang | Production of light aromatic hydrocarbons from biomass by catalytic pyrolysis[END_REF].

CaO is considered as one of the base reactants used for pyro-gasification tar upgrading during MSW thermal conversions [START_REF] Kobayashi | Woody biomass and RPF gasification using reforming catalyst and calcium oxide[END_REF][START_REF] Tang | Effect of Calcium Hydroxide on the Pyrolysis Behavior of Sewage Sludge: Reaction Characteristics and Kinetics[END_REF][START_REF] Wang | Co-pyrolysis of bamboo residual with waste tire over dual catalytic stage of CaO and co-modified HZSM-5[END_REF]. According to Kabir et al. [START_REF] Kabir | Recent progress on catalytic pyrolysis of lignocellulosic biomass to high-grade bio-oil and bio-chemicals[END_REF], the performance of CaO is quite encouraging; the addition of CaO together with the feedstock significantly changes the compositions of tar by eliminating acids and reducing phenol as compared with non-catalytic tar species; also it is effective to reduce the relative content of linear aldehydes, but promotes the formation of cycloalkanes and aromatics. Wang et al. compared the non-catalytic and catalytic pyrolysis of corncob using thermogravimetry analyzer coupled with Fourier transform infrared spectroscopy (TGA-FTIR) apparatus, and observed that the pyrolytic tar is completely deacidified by adding CaO reactant and the conversion of acid compounds promotes the generation of hydrocarbons and CH 4 ; besides, the oxygenated compounds are also decreased due to the addition of CaO reactant [START_REF] Wang | Comparison of catalytic pyrolysis of biomass with MCM-41 and CaO catalysts by using TGA-FTIR analysis[END_REF]. According to Wiggers et al., the alkali and alkaline metal oxides as base reactants can be used to reduce the oxygenated functions in bio-oil to hydrocarbons by catalytic decarboxylation and decarbonylation reactions [START_REF] Wiggers | Renewable Hydrocarbons from Triglyceride's Thermal Cracking[END_REF].

Accordingly, the role of CaO reactant on MSW pyrolytic tar reduction and tar composition upgrading are investigated in Chapter 5 and Chapter 6 in the present thesis.

Life cycle assessment and its application on MSW management

Life cycle assessment (LCA) modeling is also conducted in the present work of MSW thermal conversion system, especially pyro-gasification-based WtE systems. LCA is an evaluation methodology concerning the entire life cycle stages of a system from cradle to grave, i.e., from raw material acquisition throughout products processing, manufacture, transportation, utilization, and final disposal or recycling [START_REF] Meneses | Life cycle assessment as an environmental evaluation tool for control strategies in wastewater treatment plants[END_REF]. It is well recognized as a holistic and systematic methodology and it is effective to quantify the relevant energy and material input and output streams and the overall environmental burdens of a specified system [START_REF] Kamo | Overview of the waste plastic recycling system in Japan and future tasks[END_REF][START_REF] Lee | Assessing environmentally friendly recycling methods for composite bodies of railway rolling stock using life-cycle analysis[END_REF].

Principles and framework of LCA

LCA is firstly defined and developed by SETAC (Society of Environmental Toxicology and Chemistry) in 1990 [START_REF]Guidelines for life-cycle assessment: a code of practice: Pensacola[END_REF]. And then it is regulated by ISO standards concerning the principles and frameworks of LCA and the definition and implementation of LCA is thus standardized [START_REF]Environmental Management: Life Cycle Assessment; Principles and Framework: ISO[END_REF][121][START_REF] Klöpffer | The critical review of life cycle assessment studies according to ISO 14040 and 14044[END_REF]. A general consensus on the methodological framework between the two bodies has started to emerge, fo i g toda s fou -phase LCA framework as shown schematically in 4) interpretation [START_REF]Environmental Management: Life Cycle Assessment; Principles and Framework: ISO[END_REF][121][START_REF] Klöpffer | The critical review of life cycle assessment studies according to ISO 14040 and 14044[END_REF]. Accordingly, the implementation of LCA in waste thermal conversion systems is summarized in details as follows:

Goal and scope definition

Goal and scope definition is the most fundamental step of a LCA study, which explains the purpose of the study and determines the functional unit (FU), system boundary and allocation decisions [START_REF] Clift | The application of life cycle assessment to integrated solid waste management: Part 1-Methodology[END_REF][START_REF] Zhou | Waste-to-energy: A review of life cycle assessment and its extension methods[END_REF]. Typically for LCA studies associated to waste management, the aims at evaluate the environmental impacts caused by the implementation of waste management activities. Figure 1.10 summarizes the concerns of recent LCA studies in waste management fields [START_REF] Laurent | Review of LCA studies of solid waste management systems-Part I: Lessons learned and perspectives[END_REF], mainly including: 1) waste generation and collection; 2) waste treatment, biological treatment (composting, anaerobic digestion, bio-drying, and etc.), mechanical and chemical pretreatment, and thermal treatment (incineration, gasification, co-combustion, pyrolysis, and etc.); and 3) final use or disposal of the products. When conducting a specified case study, the techniques, life cycle stages as well as the boundary of the scenario must be defined and explained according to the aim of the study.

Life cycle inventory

Life Cycle Inventory (LCI) analysis is to create a data inventory to quantify input and output energy and material flows. The input of water, energy, raw material and the output of air emissions, waste water, solid residues and the products (energy/material) should all be included in a thorough LCA analysis. To develop the inventory, a flow model of the technical system is usually and illustrated with a flow chart that includes the activities, inputs and outputs. Data used in LCI must be related to the functional unit.

As it is well recognized, the availability and quality of inventory data are the most important part of LCA studies [START_REF] Thanh | Assessment of potential impacts of municipal solid waste treatment alternatives by using life cycle approach: a case study in Vietnam[END_REF]. However, detailed operational data for waste management systems, especially pyro-gasification-based WtE systems, are usually hard to acquire, either due to its scarce commercial applications or because those data are usually kept confidential. Fortunately, thanks to the contributions by some companies and research groups, several databases and software are developed for commercial applications to help with the establishment of LCA studies, for example Ecoinvent database, GaBi Database, SimaPro software, GaBi software, and etc. [START_REF] Martínez-Rocamora | LCA databases focused on construction materials: A review[END_REF][START_REF] Speck | Life cycle assessment software: Selection can impact results[END_REF]. In this work, LCA case studies of pyro-gasification-based WtE systems are performed using GaBi software and GaBi Database.

Life cycle impact assessment

Inventory analysis is followed by life cycle impact assessment (LCIA). The aim of LCIA is

to evaluate the significance of potential environmental impacts based on the inventory data.

Generally, a LCIA usually consist of the following steps:

-Selection of the impact assessment methods. Numerous impact assessment methods have been recently developed including: CML, Eco-indicator 99, ReCiPe, EDIP2003, and etc.

-Selection of the considered impact categories. Based on the impact assessment method selected, the impact categories should be identified, including: acidification, climate change, depletion of abiotic resources, ecotoxicity, eutrophication, human toxicity, ozone layer depletion, and so on.

-Classification of the inventory data. Based on the data inventory, emissions are classified and assigned to specific impact categories.

-Quantification of the contributions to the chosen impact categories using characterization factors. The categorized inventory data are unified to the same equivalence units by characterization factors (characterization).

-Further optional steps including normalization, grouping, and weighting may be conducted depending on the goal and scope of the LCA study.

Life cycle interpretation

The final stage of a LCA study is to interpret the results from LCI and/or LCIA to identify, quantify, check, and evaluate information of the considered systems. According to ISO 14040 [START_REF]Environmental Management: Life Cycle Assessment; Principles and Framework: ISO[END_REF], the LCA interpretation should include the following steps:

-identify and compare the environmental impacts of the contributors on basis of LCI and LCIA results;

-evaluate the systems by considering the aim defined in goal and scope definition stage;

-validate the parameters adopted by sensitivity analysis and consistency checks;

-and, draw conclusions, discussions and recommendations.

Structure and research content of the study

The structure and research content of this thesis is shown in Figure 1. 11, with a brief introduction of each chapter summarized as follows.

Background and literature review

Finally, the general conclusions related to the present work are summarized, and prospects are proposed for future research works to complete the current investigations.

Chapter 2

Materials and methods

In this chapter, feedstock and reactants used in this study are characterized; the experimental setup, working conditions, products sampling and analysis method, and life cycle assessment methodology are explained. This chapter can be divided into four parts and the contents are:

- Pure and Applied Chemistry (IUPAC) [3], hysteresis loop in multilayer range of isotherms is usually associated with capillary condensation in mesopore structures. The hysteresis loop of the CaO reactant is matched by the Type H3, which does not exhibit any limiting adsorption amount even at high P/P 0 conditions. Therefore, the pores are characterized as the aggregates of slit-shaped pores.

CaO reactant

The N 2 adsorption-desorption logarithmic isotherms curves are also plotted in Figure 2.

3.

According to Ohba [4,5], pore size distribution of the sorbent can be divided into several categories based on relative pressure of the logarithmic isotherms: adsorption in relative pressure range of 10 -6 -10 -4 is mainly associated with ultramicropore filling; adsorption in P/P 0 range of 10 -4 -10 -2 is related to monolayer adsorption of micropores; adsorption in the range of 10 -2 -10 -1 is considered to be pore filling between adsorption layers in micropores; in P/P 0 range of 0.1-0.9, the adsorption is mainly classed into monolayer-multilayer adsorption and capillary condensation in mesopores; and in the range of 0.9-1, the adsorption is mainly regarded as capillary condensation in macropores. As illustrated in Figure 2. 3, N 2 adsorption increases significantly in relative pressure range of 0.1-1, indicating that there is a great deal of mesopores and macropores existing in the selected reactant.

For each chapter of the present work, different feedstock and reactant are used. Table 2. 4 summarizes the details. The experimental system consists mainly of: 1) gas source unit; 2) electric heater and temperature controller unit; 3) main reaction chamber; and 4) HCl gas analysis unit.

-gas source unit: in this experiment, high pressure steel cylinders are used as the gas source, the used gas includes: 2000 part per million volume (ppmv) HCl gas (nitrogen balance), CO -main reaction chamber: due to the high corrosion properties of HCl gas, a vertical quartz tubular reactor is adopted, with 600.0 mm in height and 22.0 mm in inner diameter;

-HCl gas analysis unit: the off-gas is detected by Gasmet equipment (GASMET 4030;

Gasmet Technologies Oy, Pulttitie 8A, FI-00880 Helsinki, Finland); and the exhaust gas is cleaned by a bottle of 0.1 mol/L sodium hydroxide (NaOH) solution before being discharged.

Operating conditions and experimental procedure

The effects of absorption temperature, inlet HCl gas concentration, and inlet gas composition on HCl mitigation by CaO reactant are investigated in this part.

Influence of CaO reactant on HCl release and mitigation from inorganic and organic sources of chlorine 2.2.2.1 Fixed bed tubular furnace experimental set-up

Fixed bed tubular furnace experimental set-up is used to investigate the influences of temperature and CaO absorbent on HCl release and control. As depicted in Figure 2. 5, the experimental system consists of three units: 1) gas source unit, 2) electrical tubular furnace unit, and 3) sampling unit. -electric heater and temperature controller unit: a programmed temperature control heating system is used in this experiment, with the precision of the temperature controller of ±1 o C; the main reactor is a quartz tube, with 40.0 mm in inner diameter and 700.0 mm in length; feedstock is placed in a quartz boat during the experiments;

-sampling unit: tar sampling is conducted by a water-cooling condenser; the produced HCl gas is trapped by 0.1 mol/L NaOH solution; incondensable gas products are collected by gas bag.

Operating conditions and experimental procedure

The effect of temperature and addition of CaO reactant on Cl-content contaminants released from inorganic chlorine (NaCl) and organic chlorine (PVC) thermal conversion are experimentally investigated. For the experimental runs, pure N 2 is used as the carrier gas and supplied by the gas source. The tubular furnace is electrically-heated, with a thermocouple inserted in its middle to monitor the temperature profile. The quartz tube is 40.0 mm in inner diameter and 700.0 mm in length. During the experiments, the produced gas passes firstly a condenser for tar collection; and then, three impingers are used to trap the produced HCl gas. The impingers are immersed in an ice-water bath; among the impingers, the first two are filled with 0.1 mol/L NaOH solution, and the third one contains silica gel for moisture removal.

Thermogravimetric Analyzer (TGA)

A The role of CaO reactant on pyrolytic tar composition upgrading from waste single components is investigated using a Py-GC/MS experimental set-up. As illustrated in Figure 2.6, the experimental system consists of two parts: 1) pyrolysis reaction part; and 2) gas chromatography (GC) -mass spectrometry (MS) detection part. identical for all the experimental runs as 30 min. The produced gas is collected by a gas bag during this period. At the end of each test, the electrical heater is turned off immediately, but the carrier gas is kept until the temperature is cooled down to room temperature. Table 2.8 summarizes the experimental working conditions adopted in chapter 6.

The effect of temperature and CaO reactant on pyrolytic product properties (for example product mass distribution, produced gas composition, tar yield and composition, solids yield and etc.) and HCl mitigation efficiency is investigated. A temperature series of 550, 650, 750, and 850 o C is examined.

Sampling and analysis

HCl sampling and analysis

The sampling method of HCl produced from waste pyrolysis is in accordance with the USA EPA METHOD 0051. As depicted in In order to determine the effect of CaO reactant on HCl gas removal, HCl mitigation efficiency (HME) is defined as follows:

     
HCl generation with CaO reactant HME= 1-x100% HCl generation without CaO reactant (2.2) where, HRE represents for HCl mitigation efficiency.

Tar sampling and analysis

Tar yield and tar composition are investigated in Chapter 5 and Chapter 6. Tar is collected from the water cooling condenser as depicted in Figure 2.5. Liquid products in condenser as well as that adhered on wall surface and in connection pipe are washed and extracted by dichloromethane solution. Then it is separated into organic and inorganic fractions. After evaporating the dichloromethane content of the organic fraction, the remaining part is treated as tar yield. Beside the total yield, effect of temperature and CaO reactant on tar composition is investigated. Tar composition is identified by Gas Chromatography/Mass Spectrometry (GC/MS, Agilent 7890B/5977A). The percentage of the identified tar compounds is represented by the ratio of the corresponding peak area to the total peak areas.

Produced gas sampling and analysis

Produced gas from S-MSW pyrolysis (Chapter 6) is determined and analyzed to evaluate the effect of CaO reactant on produced gas quality enhancement. All the produced gas is collected in a gas bag during the whole period of each experimental run. A micro gas chromatograph (micro GC-490 analyzer, Agilent) is used to measure the molar composition of the produced gas, and the gas contains mainly N 2 , CO, CO

    2 2 2 - - N N N gas S MSW V V x Y m (2.3)
where, Y gas is the total gas yield from S-MSW pyrolysis, mL/g-S-MSW; V N2 represents the total volume of the purge gas N 2 during the experimental period, mL; x N2 is the molecular fraction of N 2 in the produced gas, %; and m S-MSW is the mass of the feedstock, g.

Solid characterization

Solid samples, including selected MSW feedstock, solid products and CaO reactant are characterized by using TGA, FESEM, XRD/EDS, XRF and BET. Details are presented as follows.

-X-ray fluorescence (XRF)

XRF analysis is performed with an Energy Dispersive X-ray fluorescence spectrometer (EPSILON 3XL) from PANanalytical company. The electrical power of the X-ray is 15 W. Filters are positioned on the beam route to the sample to avoid detector saturation. The resolution range of the high resolution Silicium Drifft SDD detector is 135 eV to 5.9 keV.

-Scanning Electron Microscope (SEM) and Energy-dispersive X-ray spectroscopy

(EDS) measurement

A scanning electron microscope (SEM, Carl Zeiss, Utral 55) is used to observe the surface morphology of both the reactants and the solid products by scanning the sample surface with a focused beam of electrons. The electrons are accelerated through a high voltage (20 kV). And the electrons interact with atoms in the sample and produce signals containing information about the s surface morphology of the sample.

EDS is used to detect the elemental composition and chemical characterization of a sample surface by determining the interaction of X-ray excitation and sample surfaces.

-X-ray diffraction (XRD)

XRD is used to identify molecular structure of a crystal by diffracting the incident -BET measurement BET measurement is used to determine the specific surface area, pore volume and structure of solid samples. Determination of the specific surface area is based on the adsorption isotherm. More detailed knowledge for the operating mechanism of BET could be found in the work from Ducousso [7]. The used BET analyzer is an ASAP 2010 type from Micromeritics. Argon is used as the adsorption gas, and the Brunauer-Emmett-Teller model is used for the determination of specific surface area; pore size and pore volume is evaluated using the Horvath-Kawazoe model. Based on the results, porosity, which represents the fraction of void space in a specified material, could be obtained by equation (2.4):

p  T V P (%) = 100% V (2.4)
where, V p stands for the total pore volume detected by BET analysis; V T is the total or bulk volume of material; P denotes porosity.

Chapter 2: Materials and methods

Mass balance and chlorine balance calculation

To make sure the correctness of the interpretation of the experimental results, experiments are repeated at least three times and the reported data are the average value of the trials. Besides, mass balance of each experiment is calculated by comparing the input flows such as feedstock, reactant, and the gas agent and the output flows like produced gas, condensate and solid residues. Mass balances of all the experimental runs are in the range of 85-103%, which is quite acceptable to ensure the quality of experimental data. Incomplete recovery of condensates, estimation of the produced gas yield and the uncertainties of the measurement may be responsible for the deviation of the mass balance closure from 100%.

Special attention has been paid to the chlorine content in feedstock as well as in products. Chlorine balance is also calculated during MSW pyrolysis experiments. In contrast to mass balance, chlorine balance closure is not achieved; chlorine balances for all the experimental runs varies from 75% to 92%, mainly due to the small amount of total chlorine content and loss of HCl in the sampling line. principle of minimizing the total Gibbs free energy to realize the thermodynamic equilibrium of the considered system. Generally, Gibbs free energy is considered as a thermodynamic potential to quantify the useful potential work, which can be acquired from the target system [1,8,9]. For a given system with constant temperature and pressure, when it reaches thermodynamic equilibrium state, the total amount of the Gibbs free energy is minimized.

Therefore, the principle of minimizing the total Gibbs free energy can be used to estimate the possible product species for a given system according to its elemental composition.

During the simulation, the system input includes the working conditions, feedstock and the reactant, and the final equilibrium state is determined by calculating all the possible species derived from the input elements with the minimum Gibbs free energy.

For a given condition of specified temperature and pressure, the total Gibbs free energy of the considered system can be obtained by:

  • 0, 1, 2,3,..., i i i G n n i N      (2.5)
Where n i represents the moles of the i th possible component in the system, u i is the free energy of the i th possible component, N is the total amount of the possible species in the considered system.

When the considered system reaches equilibrium state, the proper n i values for the components should minimize the function G. Besides, n i should also satisfy the following conditions:

1)

0 i n  ;
2) And for a given element j,

• ij i j i a n B  
, where B j is the total molar amount of element j, a ij is the molecular amount of element j in the i th component (satisfied with the fundamental principle of mass conservation).

Thermodynamic simulation software FactSage 6.3

In this thesis, thermodynamic simulation software, FactSage 6.3, is used to perform the theoretical calculations as presented in The FactSage software consists of a series of modules, including general information, database, calculation (reaction, predom, Eph, Equilib, Phase Diagram, and OptiSage) and manipulation modules [12]. Among which, the Equilib calculation module is one of the most widely used program. As aforementioned, the software uses the Gibbs energy minimization algorithm to calculate the concentrations of all possible chemical species when specified elements or compounds react or partially react to reach a chemical equilibrium state at given conditions.

In this work, the release of HCl from chlorine sources is theoretically simulated by FactSage 6.3; the effect of temperature, atmosphere, moisture content as well as reactants on product distribution is identified. The obtained simulation results provide a better understanding of the experimental results and it can be used from reaction mechanism perspective to explain the experimental results when both experimental and simulative data are assimilated.

Life cycle assessment

The definition and general introduction of LCA is explained in Chapter 1. This section details the LCIA methodology adopted in the present work. In this study, the CML (baseline) methodology (created by the University of Leiden in the Netherlands) is used [13]. The CML impact category method is one of the most popular and well-acceptable LCIA methodology [14] and the impact categories are illustrated in Table 2. 9. the presence of HCl in flue gas has already acted the role of corrosive agent and chorine source for polychlorinated dibenzo-p-dioxins and polychlorinated dibenzofurans formation during flue gas cooling down under such conditions [3]. Therefore, the primary aim of this work is to investigate the role of CaO reactant on in-furnace HCl immobilization to avoid the release of HCl into produced gases. Under such circumstance, the obtained high quality pyro-gasification products can be further utilized in an energy efficient and environmental friendly approach.

As aforementioned in Chapter 1, calcium-based absorbents are the most commonly used reactants due to their abundant sources, relatively cheap and high removal efficiency.

Therefore, CaO reactant is selected in the presented work. In order to investigate the effect of in-situ CaO reactant on HCl capture, gas-solid reaction mechanism between CaO and HCl is investigated in this chapter by performing breakthrough experiments as well as thermodynamic and kinetic modeling. Accordingly, the structure of this chapter is presented as follows:

-Section 3.2 theoretically investigates the gas-solid reaction mechanisms between CaO and HCl. The Gibbs free energy of the reaction and standard equilibrium constant is estimated by thermodynamic analysis. Kinetic modeling using shrinking unreacted core model is also performed; -Section 3.3 is dedicated to investigate the effect of key parameters on HCl mitigation by CaO reactant. The varied conditions include the reaction temperature, raw HCl gas concentration and the presence of other gas species. Morphology analysis of the spent CaO reactants is also performed to determine the formation of product-layer;

-Section 3.4 summarizes the findings obtained in this chapter.

Gas-solid reaction mechanism between CaO and HCl

Thermodynamic analysis of the reaction between CaO and HCl

The well recognized absorption reaction of HCl gas by CaO is described as follows [4,5]:

        2 2 2 s g s g CaO HCl CaCl H O    (3.1)
The absorption reaction is actually reversible at high temperature, which means temperature will affect the thermodynamic equilibrium of the reaction. Therefore, thermodynamic analysis of dehydrochlorination reaction is of great interest. According to the reaction presented in Eq. (3.1), the reaction enthalpy, Gibbs free energy and standard chemical reaction equilibrium constant of the dehydrochlorination reaction at standard temperature and pressure (STP) can be estimated by: , , products reactants

r m i r m j i r m j i i H r H r H                      (3.2) , , products reactants r m i r m j i r m j i i G r G r G                       (3.3) / rm InK G RT     (3.4)
he e, Δ r H m θ a d Δ r G m θ are enthalpy of formation and Gibbs free energies of formation at STP, respectively; j refers to a specific product or reactant and r i refers to its stoichiometric coefficient; K θ is standard chemical reaction equilibrium constant at STP and R is ideal gas constant and the value is 8.314 J/(K*mol). However, considering the high reaction te pe atu e i this o k, the effe t of te pe atu e o Δ r H m θ , Δ r G m θ , and K θ must be taken into account. The Gibbs-Helmholtz equation reveals that the change of Gibbs free energy of a system is a function of temperature as [6]:

  2 /T P G H T T           (3.5)
Therefore, for reactions at constant pressure, Eq.(3.5) becomes:

  2 /T rm rm G H T T      (3.6)
Combining Eq.(3.6) and the integral form of Eq.(3.4), the reaction constant as a function of temperature can be obtained as:

    0 0 2 T rm T H InK T InK T dT RT      (3.7)
A o di g to the Ki hhoff s la of thermochemistry, Δ r H m can be obtained as [7]:

    0 0 p d T r m r m r T H T H T C T        (3.8)
where, C p represents the specific heat capacity at constant pressure, and it can be estimated by the Shomate Equation according to [8,9]:

23 p 2 e C a b T c T d T T         (3.9)
A d Δ r H m can thus be obtained by [8]:

    2 3 4 0 /2 /3 /4 / r m r m H T H T a T b T c T d T e T f h                       (3.10)
Where a, b, c, d, e, f, and h are the coefficients for the calculation of the specific heat capacity at temperature of T. A o di gl , he Δ r H m is determined, K θ can be calculated based on Eq. (3.7); a d si ila l , Δ r H m under different temperature can be obtained based on Eq. (3.4). Table 3.1 summarizes the specific heat capacity coefficients and thermodynamic data for the reactants and products from Eq.(3.1) [10]. The calculation results are reported in Table 3.2, results reveal that within the temperature range of 500-1200 K, the Gibbs free energy of the reaction is always negative, indicating that the gas-solid reaction between CaO and HCl could spontaneously occur.

However, with the increasing of temperature, the Gibbs free energy of reaction rises continuously and the standard equilibrium constant declines significantly by approximately twelve orders of magnitude from 6.97X10 16 at temperature of 500K to 2.28 X10 4 at temperature of 1200K. This means that although the reaction is still spontaneous, the degree of the reaction toward the formation of CaCl 2 and H 2 O will significantly be inhibited due to the shift of chemical equilibrium. It is notable that although the reaction is spontaneous from thermodynamic perspective, the reaction rate is significantly affected by the kinetics; therefore, kinetic investigation of the considered reaction system is of great importance.

Kinetics analysis of the reaction between CaO and HCl

As aforementioned in last section, thermodynamic analysis determines whether the reaction system is spontaneous; but the reaction rate is affected by the reaction kinetics.

Therefore, kinetics modeling of the considered reaction system (as illustrated in Eq. (3.1)) is investigated in this section.

The grain model and shrinking unreacted core model are often used to describe a non-catalytic gas-solid reaction for porous and nonporous solid reactant, respectively, with the purpose to determine the rate-limiting step of the process by either chemical reaction or product-layer diffusion (PLD) and to estimate the key kinetic parameters, for example chemical reaction rate constant and diffusivity for PLD process [4,11]. The shrinking unreacted core model is used in this study, because the used CaO reactant is nonporous [12]. product-layer is possible, since it is considered porous. Also, it is assumed that the surface reaction is considered so fast that as soon as HCl reaches the surface of the unreacted CaO core, the reaction occurs and HCl is consumed. Therefore, the reaction surface (i.e. the core surface) moves towards the center of spherical CaO particle. However, as presented in Figure 3.1, although the formed product-layer is mainly consisted of CaCl 2 , it is assumed in the simplified theoretical model that the overall particle size (R, represents the overall particle size of both the product-layer and the unreacted CaO core) remains unchanged.

Ideal shrinking unreacted core model

The reaction between the isothermal spherical CaO particle and a constant concentration of HCl gas flow (from the ambient environment) can be described as [13,14]:

-external diffusion step: diffusion of the HCl gas through the surrounding gas film of the solid particle to reach its surface;

-internal diffusion step: diffusion of the HCl gas through the porous product-layer (the produced CaCl 2 layer considered in this work) to the unreacted CaO surface;

-chemical reaction step: chemical reaction occurs between HCl gas and unreacted solid CaO on the core surface;

-internal diffusion step: diffusion of the product gas H 2 O vapor through the porous product-layer to the solid surface;

-And external diffusion: diffusion of the product gas H 2 O through the gas film to the bulk. Under such circumstance, the conversion rate of CaO can be described by [4]:

As depicted in

  3 1- = = c CaO g r t X G x R      (3.13)
where, X CaO is the conversion rate of solid reactant CaO; k g represents the mass transfer constant of HCl gas surrounding the reactant; a d τ g is the time needed for the total conversion of CaO a d τ g = 6 Under such condition, the conversion rate of CaO can be described by [15]:

    2 3 2 3 1 3 2 1 3 1 2 1 () c c CaO CaO p r r t R R X X Px                             (3.

Chapter 3: Experimental and kinetics study on HCl mitigation by CaO reactant

Under such condition, the conversion rate of CaO can be described by [15]: linearly dependent on the reaction time t, the overall reaction is considered to be controlled by external diffusion; if P(x) is linearly dependent on the reaction time t, the overall reaction is considered to be controlled by internal diffusion; and if C(x) is linearly dependent on the reaction time t, the overall reaction is considered to be controlled by chemical reaction rate.

  1 3 1 1 1 ( ) c CaO c r t X C x R                 (3.

Estimation of kinetic parameters

Previous studies revealed that compared with internal diffusion and chemical reaction rate, external diffusion of gas-solid reactions, i.e. the resistance of reactant gas from the bulk to the solid particle, is not significant and is negligible [16][17][18]. Therefore, it is not taken into account in this work.

When considering the change of the particle size, the conversion rate of CaO under the condition that the internal diffusion is rate-limiting can be modified to the following equation [4,15]:

       2/3 2/3 11 ( ) 3 1 1 CaO CaO Z Z Z X P x X Z                   (3.16)
where, Z represents the ratio of molar volume of product CaCl 2 to that of reactant CaO, and the value is selected as 3.04 in the present work.

Therefore, the reaction time t and the conversion rate of CaO reactant can be described when both internal diffusion of HCl in the product-layer and chemical reaction on CaO core surface are the rate-limiting steps:

2 ( ) ( ) c t C x P x    (3.17)
where,

2 6 p s ce kR D    
represents the ratio of the product-layer diffusion resistance to the chemical reaction resistance, which can be used to determine whether the process is limited by diffusion or chemical reaction rate. When 2  <<1, it means the diffusion resistance is much lower than the reaction resistance and the process is controlled by chemical reaction; when 2  >10, the process is mainly governed by product-layer diffusion;

and when 1< 2  <10, it means both product-layer diffusion and chemical reaction are determining factors for the process.

By least-squares fitting of Eq. (3.17) to the e pe i e tal data, alues of τ p , τ c , and 2  can be determined. And then the effective diffusivity of HCl through the porous CaCl 2 product-layer D e and the chemical reaction rate k s can be calculated. According to the Arrhenius equation, the chemical reaction rate k s can be determined by:

exp s E kA RT      (3.18)
where A is the pre-exponential factor in min -1 ; E is the activation energy in kJ/mol; R is the ideal gas constant and its value is 8.314 J/(mol*K). And therefore:

ln ( / )(1/ ) s C k E R T  (3.19) ln ( / )(1/ ) e P D E R T  (3.20)
Where, E c and E p represent the apparent activation energy of chemical reaction and product-layer diffusion, respectively.

Experimental results

Mitigation of HCl gas by CaO reactant under moderate high temperature are investigated in a vertical tubular reactor and the experimental set-up, working conditions and experimental procedures are described in Chapter 2. The effect of reaction temperature, raw HCl gas concentration, purge gas velocity, and purge gas composition are determined and the kinetic parameters are estimated according to experimental data. In this section, several parameters to explain the experimental results are defined as:

Breakthrough time: the time when the outlet concentration of HCl is detected to be higher than 20 ppm, (due to the fluctuation of the detector, 20 ppm (i.e. c/c 0 = 20 ppm /2000 ppm = 0.01) instead of 1 ppm is selected as the limit for breakthrough; in cases of varying the raw HCl concentration, 10 ppm for c 0 = 1000 ppm and 5 ppm for c 0 = 500 ppm is used to measure the breakthrough time); Saturation time: the time when the outlet concentration of HCl is detected to be higher

tt t t HCl HCl CaO HCl c c V dt m V m k t t            (3.22)
where, k HCl represents the HCl mitigation ate; Δ HCl is the absorption amount of

HCl o e a pe iod of ti e Δt. The results are reported as follows:

Figure 3. 6 The HCl breakthrough curves of CaO reactants under different temperature

It is well recognized that temperature is a dominant factor determining gas-solid absorption process by affecting both physical adsorption and chemical reaction. On the one hand, from thermodynamic point of view, physical adsorption is an exothermic process, therefore, increasing the temperature will cause adverse effects on physical adsorption of gases by solid sorbents; and also, chemical reaction highly relies on temperature by breaking chemical bonds as well as shifting the chemical equilibrium. On the other hand, kinetically speaking, higher temperature will enhances the diffusion rate of HCl in the product-layer, but at higher temperature, the surface properties (for example specific surface area, pore structure and pore volume) of the sorbent will significantly be changed, thus the diffusion of gases and chemical reaction rate on particle surface will be influenced. Results from Section 3.1 demonstrated that higher temperature will shift the dehydrochlorination reaction equilibrium towards the direction of CaCl 2 decomposition. Therefore, increasing the According to Section 3.2.2.2, the corresponding kinetic paramete s τ p , τ c , θ 2 , k s , D p , E c and E p are listed in Table 3.3. A o di g to the esti atio esults, the alue of θ 2 is always in the a ge of < θ 2 < 10, which indicates that the absorption of HCl by CaO reactants is limited by both chemical reaction and product-layer diffusion. The phenomena further confirm that the reduction of surface area and closure of pore structures are responsible for the decline of HCl mitigation efficiency at higher temperature. On the contrary, the specific surface area and pore volume are slightly increased to 2.1 m 2 /g and 2.1x10 -2 to 1.6x10 -2 cm 3 /g at temperature of 850 o C. This is mainly due to the decomposition and evaporation of gas-solid reaction products (i.e. the product-layer). However, at temperature of 850 o C, HCl mitigation efficiency is also reduced as compared with that at 750 o C, which indicates that the shift of chemical reaction equilibrium plays a more important role under such circumstance.

Effect of raw HCl concentration on HCl mitigation by CaO reactant

Effect of carrier gas composition on HCl mitigation by CaO reactant

Previous concerns are mainly focused on the absorption of HCl in N 2 atmosphere; however, for actual waste pyro-gasification systems, the presence of HCl in the produced gas is always accompanied with other gases compositions, which may also react with CaO reactant and thus influence the HCl mitigation efficiency. Therefore, the effect of gas The effect of gas compositions on total HCl mitigation amount and the mitigation rate are presented in

Summary of the chapter

In this chapter, the gas-solid reaction mechanism between CaO reactant and HCl gas is studied by both theoretical calculation and experimental investigation. Thermodynamic analysis is firstly conducted to determine the Gibbs free energy and standard equilibrium constant of the reaction. Then breakthrough experiments and kinetic modeling of the reaction are performed to identify the effect of key parameters on HCl mitigation efficiency.

Last but not least, the micrograph of the spent CaO reactant is examined to reveal the formation mechanism of the product-layer.

The following conclusions can be obtained:

 Thermodynamic analysis reveals that the reaction between CaO and HCl is inhibited when increasing the temperature due to the shift of chemical reaction equilibrium.

 Experimental results reveal that increasing the temperature, reducing raw concentration of HCl gas, and the presence of CO, and H 2 in carrier gas cause the reduction of total HCl mitigation amount and CaO bed conversion efficiency. 12.4% is employed to produce inorganics, and the rest is used for the generation of other organics [1]. Considering the important role of chlorine in modern chemical industry as well as the abundant amount of inorganic chlorides existing in the natural world, chlorine content in MSW is as high as 800-2500 mg/kg according to Jhong-Lin Wu et al [2].

As aforementioned in Section 1.2, the presence of chlorine in MSW is mainly in two forms: in inorganic form, such as NaCl or KCl; and in organic chlorine source form like PVC or rubbers [3][4][5]. During waste thermal treatment, chlorine is liable to be released as a

Chapter 4: Influence of CaO reactant on HCl release and control from inorganic and organic sources of chlorine 90 consequence of high temperature. And the produced Cl-compounds may participate in various reactions in the furnace, leading to several serious consequences, for example, high temperature corrosion, heavy metal evaporation and dioxins and furans formation. Besides, the emitted Cl-content species, especially toxic organic compounds, are considered to be persistent and have the potential to migrate widely throughout the ecosystem [6,7].

Therefore, the formation of chlorine-contained contaminants from different kinds of chlorine sources is of great importance. The primary aim of this chapter is to investigate the fate of chlorine from both inorganic and organic sources.

Chlorinated pollutants release from inorganic and organic sources 4.2.1 Chlorinated pollutants from inorganic source-NaCl

The most typically inorganic chlorine source existing in food waste, NaCl, is used as the feedstock for experimental purposes. FactSage software is used to carry out the thermodynamic equilibrium simulation. Experiments are conducted on both thermogravimetric and fixed bed tubular furnace apparatus.

Thermodynamic equilibrium simulation of Cl-contained contaminants from NaCl

Thermodynamic simulations of the thermal behavior of NaCl under the considered working conditions are firstly theoretically investigated using FactSage software. The results could be served as the fundamental information to explain the phenomenon observed during the experiments. Simulation is performed using the same input condition of the experimental studies. (NaCl) 2 (g) in the total Cl-content species reach 0.45% and 0.33%, respectively. The molar fraction of NaCl(liq) decreases to around 99.22%.

The effect of water content on the thermodynamic behavior of NaCl is presented in Meanwhile, NaOH(g) is also formed according to the calculation, indicating that the chemical reaction Eq.(4.1) occurred in the equilibrium system. At temperature of 950 o C, NaCl(g), (NaCl) 2 (g) and HCl(g) in the total Cl-components are equal to 2.23%, 1.66%, and 0.05% respectively; while the molar fraction of NaCl(liq) in the total Cl-content species decreases to 96.06%. Besides, the molar fraction of sodium contained in NaOH (g) to the total Na-content species is also determined, and the amount is about 0.05%. However, both the evaporation in the forms of NaCl(g) and (NaCl) 2 (g) and the formation of HCl(g) and NaOH(g) are slightly promoted under O 2 atmosphere with the presence of water.

At temperature of 950 o C, the molar fraction of NaCl(g), (NaCl) 2 (g) and HCl(g) in the total Cl-components is 2.67%, 1.99%, and 0.06% respectively; with regard to NaCl(liq), the proportion reduces to 95.27%. The molar fraction of NaOH (g) in the total Na-components is around 0.06%.

With respect to the effect of inorganic material, the presence of SiO 2 is theoretically calculated to represent the effect of ash content on the thermodynamic behavior of chlorine from inorganic chlorine. The results are shown in By adding SiO 2 into the system, the evaporation of NaCl in the forms of NaCl(g) and (NaCl) would increase both the evaporation (in the forms of (NaCl(g) and (NaCl) 2 (g)) and the thermochemical conversion from NaCl to HCl. Given the fact that the physical composition of MSW is quite complex, the presence of water and inorganic material (SiO 2 , Al 2 O 3 , and etc.) is unavoidable; therefore, HCl release from inorganic chlorine source, for example NaCl, is worthy of further investigation. These results agree well with the observations from the thermodynamic equilibrium simulation. However, a slight mass loss of around 0.1% is observed when the temperature is higher than 750 o C, mainly due to the entrainment by the purge gas during TG experiments.

Experimental study of Cl-contained contaminants from NaCl

Even at a temperature of 850 o C, the mass loss of NaCl is around 5.2%. This fact indicates that the total decomposition or evaporation of pure NaCl is quite low in the considered temperature ranges during pyrolysis. As aforementioned, the binding energy of NaCl is as high as 787 kJ/mol; therefore, thermal decomposition or evaporation of NaCl under pyrolysis conditions is low, which has been proven in both thermodynamic simulations and TGA experiments. As illustrated in However, although a rising trend is obtained with temperature, the conversion rate is less than 1.4% even at higher temperature of 950 °C. Despite the relatively low conversion rate, it is worth mentioning that there is a remarkable increase of the rate between 750 °C and 850 °C. It is mainly due to the melting point of NaCl which is around 801°C. The conversion rate is increased by 0.6% from 750 °C to 850 °C, while the value is only 0.2% from 550 °C to 750 °C. and water will promote the release of Cl-compounds from the furnace by either evaporation or chemical conversion. Although the release level of Cl-contained contaminants from inorganic chlorine is as low as 2-8%, the remarkable amount of inorganic chlorine contained in MSW (as previously mentioned in Section 1.2.1, around 45% of the total chlorine content in MSW comes from food waste in inorganic forms) should be taken into consideration.

Besides, the high proportion of water and inorganic materials contained in MSW in some

Countries or regions will definitely facilitates the release of chlorine from inorganic source during waste thermal treatments. Therefore, inorganic chlorine may contribute to a considerable amount over the total HCl release contained in flue gas.

Chlorinated pollutants from organic source-PVC 4.2.2.1 TG and TG-FTIR analysis of Cl contaminants release from PVC pyrolysis

In order to investigate the thermal behavior of PVC during pyrolysis, thermogravimetric 3.0 a T 0.1 , T 0.9 , and T max represents for the corresponding temperature when the mass conversion rate of PVC reaches 0.1, 0.9 and the maximum value; b t 0.1-0.9 is the time duration of PVC mass conversion rate from 0.1 to 0.9; c (d a /d t ) max and (d a /d t ) mean denote the maximum and average mass loss rate during thermal conversion. formed after the dehydrochlorination process and the formation of benzene is thus enhanced [13,14].

To sum up, thermal degradation of PVC is characterized as a two-step process, in which:

-the first degradation step is in the temperature range of around 200 to 400 o C with a high mass loss percentage of 60%; dehydrochlorination reactions are mainly taking place in this stage, accompanied with the formation of a small amount of benzene;

-the second degradation step occurs in the temperature range of 400-550 o C; in this stage, further decomposition and cracking of the polymer occurs to produce light hydrocarbons due to the increase of temperature; besides, aromatic compounds, such as benzene and its derivatives are also formed due to the abundant conjugated double bonds. PVC is pyrolysed, HCl is the main gas released and the chlorine content for PVC pyrolytic char is less than 0.01 wt.%, indicating a nearly complete conversion of chlorine from PVC to HCl [15]. It is also notable that air atmosphere slightly shifts HCl release temperature to a lower level.

Influence of in-furnace CaO reactant on HCl mitigation efficiency

Because of the fact that HCl release from PVC is easy to occur in low temperature Brooke Shemwell et al [18], the binding capacity of CaO and HCl is limited at higher temperature due to the shift of the thermodynamic equilibrium of dechlorination reactions.

The evolution of the surface morphologies of the spent CaO reactant as well as their elemental composition analysis with the increase of temperature is depicted in The elemental compositions of the selected surface of the sorbent are also depicted in Results reveal that the particle size of the reactant also plays a quite important role in HCl mitigation efficiency. Higher HCl mitigation efficiency is obtained when finer CaO reactant is used in the furnace under the experimental situations. As presented in Figure 4. The specific surface area of the CaO with the particle size of 80-120 mesh is 6.28 m 2 /g, nearly 2.5 times higher than that of the CaO of 60-80 mesh (2.37 m 2 /g). Therefore, finer reactants provide more surface area (or in other words, more active sites) for the both adsorption and chemical reactions of HCl by CaO, thus can significantly enhances the HCl mitigation efficiency. From this point of view, investigation in Section 4.3.1, when the Ca/Cl molar ratio is increased, increases not only the quantity of CaO reactant, but also the number of the active sites on its surface for both HCl physical adsorption and chemical dechloirination reaction. However, it is worth mentioning that the finer the particle size, the severer the attrition of the reactants may occur, resulting in the formation of more particulates and blocking the facilities [19]. Therefore, the particle size of the reactant should be carefully optimized with regards to both the effects on HCl mitigation and the adverse effect on particulates formation especially in commercial applications.

Effect of CaO particle size on HCl mitigation by in-furnace CaO reactant

kinetic modeling of HCl release and control from PVC pyrolysis

Kinetic modeling of HCl release from PVC pyrolysis and the effect of CaO addition are investigated in this section. The model-free method is one of the most commonly used computation methods to estimate the reaction kinetic parameters (i.e. apparent activation energy and pre-exponential factor) without considering the pyrolytic mechanism. However, Chapter 4: Influence of CaO reactant on HCl release and control from inorganic and organic sources of chlorine 111 the model-free method is only suitable for a single reaction [20,21]. As aforementioned in process accompanied with a small amount of hydrocarbons, mainly aromatic compounds such as benzene [22]. Moreover, the reaction rate curve (DTG) has multiple peaks and shoulders, indicating that the pyrolytic reaction mechanisms are associated with a multi-step conversion. Regarding this fact, the model-free method is not suitable to identify the mechanisms, since it only considers the overlap of the multi-step decomposition processes [20,23].

One possible solution to the aforementioned problem is to separate the DTG curves (the reaction rate) into the combination of several individual reaction process, and then uses the model-free methods to estimate the corresponding parameters of the separated pseudo-component reactions [20]. Recently, the Fraser-Suzuki function (FSF) has successfully been applied to curve-fit complex solid-state reactions into individual reactions [20,[24][25][26][27][28].

Therefore, this method is introduced in this work.

Kinetic modeling approach 4.4.1.1 Individual reactions of PVC pyrolysis with/without CaO reactant

As aforementioned, PVC pyrolysis can be divided into three single step reactions including: FSF 1 : dehydrochlorination (de-HCl); FSF 2 : formation of aromatic compounds like benzene (de-benzene); and FSF 3 : further polyene decomposition and cyclization (de-polyene and cyclization). Therefore, it is assumed that the decomposition of the pseudo-components is independent and can be described by equations 4.9 -4.11:

  where k refers to the reaction rate constant. When CaO reactant is used, it is considered that only the reaction rate and reaction degree of the independent single reaction is influenced, but the reaction process is not changed. Therefore, the pseudo-components decomposition can be described by equation 4.12-4.14: he e k efe s to the ea tio ate o sta t ith CaO reactant.

Reaction kinetics

The kinetic expression for the rate of heterogeneous reactions can be described as:

    da k T f a dt  (4.15)
where the conversion rate a is defined as the instant mass loss to the overall mass loss, and it is given by the following equations:

0 0 t f mm a mm    (4. 16
)
where a is the conversion rate; m 0 , m t and m f represents for the initial weight, instant weight of the sample at t time and the final weight of the sample. According to the Arrhenius equation, the chemical reaction rate k(T) can be determined by:

  exp E k T A RT      (4.17)
where A is the pre-exponential factor in min -1 ; E is the activation energy in kJ/mol; R is the ideal gas constant and its value is 8.314 J/(mol*K). By combining the Eq. (4.15) and Eq.

(4.17), a d taki g i to o side atio the o sta t heati g ate, β= dT/dt, the ki eti s expressions can be described as:

  exp da A E fa dT RT          (4.18)
The integral form of Eq. (4.18) is:

    0 0 1 exp a T T A E g a da dT f a RT         (4.19)

Fraser-Suzuki deconvolution

Deconvolution of the DTG curves using the FSF function can be described as:

  2 c , , , 2 i=1
, , d ln 2 exp ln 1 2

N pi i p i s i hf i si TT a c H A dT W A                        (4.20)
where the parameters H p , T p , W hf , and A s represents the height (K -1 ), peak temperature (K), half-width of the separated DTG curves of the pseudo-components, and asymmetry (dimensionless), respectively; ci and Nc refer to the proportion of i-th pseudo-component

Activation energy determination

Several mathematic method are introduced to solve the kinetic equations and to estimate the activation energy. Among them, Ozawa-Flynn-Wall (OFW) method and Kissinger-Akahira-Sunose (KAS) method are most commonly applied [29,30].

By OFW method, Eq. (4.19) can be linearized as [13]:

    ln ln 5.331 1.052 * AE E Rg a RT         (4.21)
And by KAS method, Eq. (4.19) can be linearized as [START_REF] Coda | Behavior of chlorine and enrichment of risky elements in bubbling fluidized bed combustion of biomass and waste assisted by additives[END_REF]32]:

  2 ln ln * AR E E g a RT T          (4.22)
The efo e, the ki eti pa a ete s a e o tai ed plotti g l β s /T fo OFW ethod a d l β/T 2 ) vs 1/T for KAS method at the same conversion rate a. For the OFW Chapter 4: Influence of CaO reactant on HCl release and control from inorganic and organic sources of chlorine 114 method, the apparent activation energy can be obtained from the slope of the straight line and is equal to -1.052*E/R, and for the KAS method, the apparent activation energy can be obtained from the slope of the straight line and is equal to -E/R.

Kinetic modeling results

Fraser-Suzuki deconvolution results

The Fraser-Suzuki deconvolution curves for pure PVC and PVC with CaO reactant during pyrolysis at 5, 10 and 20 °C/min are presented in

Summary of the chapter

In this chapter, the generation of Cl-content contaminants from both inorganic and organic chlorine source is firstly investigated using both thermodynamic equilibrium simulation and experimental study. Then the effect of in-furnace CaO reactant on HCl mitigation is evaluated. Furthermore, the reaction kinetics of PVC pyrolysis without/with CaO reactant is estimated and the apparent activation energies of the pseudo-components are determined. The purpose of the study is three-folds: (i) to determine the influence of temperature, atmosphere, and the presence of water, inorganic materials on HCl release; (ii) to investigate the effect of temperature, Ca/Cl molar ratio, and CaO reactant particle size on HCl mitigation efficiency; and (iii) to identify the kinetic pyrolytic reaction mechanism of PVC with/without CaO reactant and to estimate the activation energy of the pseudo-component reactions.

The following conclusions have been drawn:



Chapter 5

Influence of CaO reactant on pyrolysis tar upgrading of waste single-component

Introduction

As it is well-known, tar is a typical by-product of pyro-gasification process, and its composition is very complex [1]. Generally, tar produced during pyro-gasification process causes syngas degradation and has been identified as one of the major impurities [2].

Moreover, the presence of tar also brings a series of problems like clogging and condensation followed by blocking of downstream equipment [3]. Therefore, tar reduction from pyro-gasification process has become one of the most popular topics in recent years [4][5][6].

Besides been treated as impurities, tar contains a large amount of high value added compounds, like benzene, toluene, and xylene; therefore, the recovery of these chemicals from pyro-gasification tar also attracts worldwide concerns.

Tar reduction can be achieved by using reactants. As presented in Figure 5.1, tar cracking and upgrading during pyrolysis can be divided into ex-situ and in-situ methods [7].

The addition of reactant into the furnace is an effective method to reduce the tar contained in the produced gas. Ca-based reactants are well recognized for pyrolytic tar cracking and upgrading [7][8][9]. Therefore, besides capturing HCl, the use of in-furnace CaO reactant is believed a promising method to accelerate the decomposition and upgrading of tar compounds during MSW pyrolysis.

Py-GC/MS is widely used as an analytical tool to determine the composition of pyrolytic tar due to its rapidity, high sensitivity, and effective separation ability of complex mixture of tar compounds [10][11][12][13]. Therefore, in this chapter, the role of CaO reactant and the effect of

Effect of CaO reactant on pyrolysis tar composition from waste single component

MSW single components pyrolysis without/with CaO reactant are conducted and investigated in this section using Py-GC/MS experimental apparatus. Pyrolytic tar compositions from experiments are identified by matching the mass spectrum with standard compounds in NIST 11 database. The relative content of each identified compound is normalized to Peak area%, which is determined by area normalizations. The relative content of a specific compound refers to the peak area of the component to the total peak area of all detected species. The calculation method of the Peak area % is described as follows:

arg % 100% det

Peak area of the t et compound Peak area

Total peak area of all ected species  

(5.1)

. Effe t of CaO reactant on tar formation from wood biomass

Wood biomass pyrolysis at temperature of 750 o C are firstly investigated using and application [14,15]. Therefore, deoxygenation is of great importance for pyrolytic tar quality upgrading and obviously, CaO can serve as deoxygenation reactant during biomass pyrolysis.

With respect to the classified groups, pyrolytic tar compounds contain mainly benzenes (accounting for 14.8% and 17.3% for non-catalytic and catalytic pyrolysis conditions, respectively) and phenols (accounting for 30.4% and 26.2% for non-catalytic and catalytic cases). The addition of CaO increases the relative content of benzenes, naphthalenes, ketones, and carbohydrates by around 17.5%, 134.5%, 173.2%, and 81.4%, respectively.

Chapter 5: Influence of CaO reactant on pyrolysis tar upgrading of waste single-component

Although the relative content of naphthalenes and ketones are dramatically enhanced (by more than 2.5 times), their relative content is quite low and in the ranges of 1.5% -3.6% and 3.6% -10.0%, respectively. The results are in good accordance with previous findings. Zhu et al. [16] found that the relative content of naphthalene and substituted naphthalenes are increased with the presence of CaO reactant. Lu et al. [8] stated that the use of CaO increases the relative content of ketones from 3.8% to 20.9%.

On the contrary, the relative contents of aldehydes, phenols, acid compounds, and heterocyclics compounds are decreased by 39.6%, 13.7%, 100.0%, and 24.9%, respectively. It is notable that the abovementioned four components are containing oxygen; therefore, the overall oxygenated compounds content is declined by adding CaO reactant. Among the oxygenated compounds, the addition of CaO reactant achieved elimination of acid compounds contained in wood biomass pyrolytic tar. It is quite logical, as it is obvious that CaO reactant is a typical alkaline earth metal oxide reactant and is effective to reduce acid compounds in pyrolytic tar by decarboxylation and decarbonylation reactions according to Kabir et al. [7]. Similar findings are reported by literature revealing that the pyrolytic tar is de-acidified by using CaO reactant [17]. As aforementioned in Section 5.1, pyrolytic tars contain a variety of high value added chemicals, mainly light aromatic hydrocarbons such as benzene, toluene, xylene and naphthalene (BTXN) [18,19]. It can further be extracted from pyrolytic tar and applied as fuel additives, solvents, and as raw material to produce pesticides, plastics, and fibers [20]. positive effects on toluene, xylene, and naphthalene formation are reported and the influence of which was not obvious on the formation of benzene [21]. As for the catalytic influence on phenols, the relative content of total phenol and substituted phenols are slightly decreased from 30.4% to 26.2% of pyrolytic tar from wood biomass; meanwhile, the relative content of phenol is dramatically reduced by around 57.4% with the presence of CaO reactant. Figure 5.6 shows a typical reaction pathway for decarboxylation and decarbonylation of pyrolytic tar from complex heavy compounds to light aromatic compounds over CaO reactant [7]. In regard to the classified groups, the content of aldehydes takes the highest proportion for pyrolysis in the absence of CaO, followed by ketones, furans, benzenes, nitriles, olefins, phenols and so on. However, with the addition of CaO, the relative content of ketones increased to 33.1%, ranking the first of all the groups, followed by benzenes, olefins, phenols, and so on. Accordingly, the addition of CaO reactant increases the relative content of olefins by 50.7%, benzenes by 181.1%, ketones by 41.3%, phenols by 39.1%, but reduces the content of aldehydes by 86.4%, furans by 86.0%, amines by 66.3%, nitriles by 27.4%. The relative content of acid compounds is quite stable at 3.8% and 3.7% under non-catalytic and catalytic conditions. The variation of the pyrolytic group compounds agree well with the results obtained from wood biomass pyrolysis, except for phenols. The relative content of phenols decreased during wood pyrolysis, but has been promoted during food waste pyrolysis when CaO reactant is used. However, the overall deoxygenation (i. 5.2. As illustrated, most of the detected compounds are aromatic compounds, and their relative content accounts for 89.5% and 93.5% of the total detected tar species from non-catalytic and catalytic pyrolysis experiments, respectively. This is mainly due to the molecular structure and the corresponding dehydrochlorination mechanism of PVC as presented in Figure 5.11 [22].

. Effe t of CaO reactant on tar formation from food waste

Results from Chapter 4.2.2 reveal that the first decomposition stage of PVC produces a large amount of HCl, leading to the formation of abundant conjugated double bond. Meanwhile, TG-FTIR experiments confirmed that dehydrochlorination of PVC is always accompanied with the formation of aromatic compounds, mainly benzenes, and the further increase of temperature will result in the acceleration of carbon chain decomposition and hydrocarbons cycloaromatization. Therefore, aromatic compounds are the dominant products during PVC pyrolysis. Different from the previous situations, no oxygenated compounds are formed from pyrolytic tar of PVC pyrolysis without CaO reactant; however, when CaO is added, a small amount of oxygen-content species are generated. It is mainly due to the fact that no oxygen is available in PVC as reported in Table 2.1; however, when CaO is added, oxygen is also fed into the system and a small amount of oxygenated compounds (CO 2 , 2.45% and 2-Hexynoic acid, 0.99%) is produced. o C. Higher temperature accelerates the cycloaromatization reactions of the generated tar compounds and thus remarkable increase the relative content of aromatic compounds [23,24]. However, for CaO catalytic pyrolysis conditions, the relative content of aromatic compounds is fluctuant in the range of 89.0% -93.3%, a little bit higher than that in the absence of CaO reactant. This agrees well with the aforementioned results that the use of CaO reactant promotes the formation of aromatic compounds. promoted. This is a counter-intuitive finding, since it is well recognized that higher temperature facilitates the thermal cracking of heavy tar compounds [25][26][27]. However, it is believed that the obtained results are reasonable. It is mainly due to the aforementioned fact that only volatile organic compounds are able to be detected from Py-GC/MS experiments, which means heavy compounds (for example tar components with more than 20 carbon atoms) cannot be recorded from Py-GC/MS apparatus. The possible fact is that when the pyrolytic temperature is increased from 550 to 850 o C, heavy tar compounds (which cannot be detected by Py-GC/MS) are thermally cracked and reduced, leading to the higher relative content of C11 -C15 compounds as identified in this work. The same phenomena are concluded by Lu et al. [28] and Zhang et al. [29], observing that the total peak area was increased monotonically and dramatically when increasing the pyrolysis temperature. It is worth noting that HCl is detected from non-catalytic pyrolysis, while it is eliminated from CaO catalytic pyrolysis under the considered temperature range. naphthalene with CaO reactant [30]. As presented, CaO acted an important role in naphthalene-content species decomposition and naphthalene formation. On the one hand, the presence of CaO will destroy the branched chains on naphthalenes, leading to the formation of naphthalene. On the other hand, at high temperature, a small amount of formed toluene loses a hydrogen atom and become free radicals; and then the free radicals react with light molecules to form naphthalene.

Summary of the chapter

In this chapter, pyrolytic tar from waste single components (wood biomass, food waste and PVC) is investigated. The role of CaO reactant and the effect of temperature on tar composition are examined.

The following conclusions can be drawn:



Chapter 6

Influence of CaO reactant on pyrolysis of simulated MSW: HCl mitigation and tar upgrading

Introduction

Apart from incineration, MSW pyrolysis is believed one of the most promising alternative WtE strategies achieving both waste treatment and energy recovery. Pyrolysis is also the most fundamental step of waste thermal conversions and determines directly the product yields and properties at the very beginning of waste thermal conversion [1].

Therefore, MSW pyrolysis is experimentally investigated in this work with the emphasis on effect of temperature and CaO reactant on both in-furnace pyrolytic product properties enhancement and contaminants (mainly tar and HCl gas) removal.

Recently, MSW pyrolysis has become one of the most popular hotspots in waste treatment fields; studies have demonstrated the effects of furnaces [2][3][4], pyrolysis temperature [5,6], pyrolysis pressure [7,8], reaction time [6,9], and heating method [10,11] on pyrolytic product properties. In addition to physical/chemical properties of MSW and the pyrolysis working condition, reactants are also a key factor affecting pyrolysis behavior [12,13]. CaO reactant also plays an important role on pyrolysis products quality enhancement.

Therefore, besides the aforementioned HCl mitigation and tar cracking and upgrading functions in Chapter 3, 4, and 5, CaO is also commonly used as the in-furnace reactant to increase the quality of the pyro-gasification products, since it can be regarded as an efficient CO 2 sorbent to increase the produced gas quality [14][15][16][17]. CaO is one of the most preferable Chapter 6: Influence of CaO reactant on pyrolysis of simulated MSW: HCl mitigation and tar upgrading 156 reactants in waste thermal conversion systems, because it is quite cheap and widely abundant on the earth [12,18,19].

Therefore, in this chapter, effects of temperature and CaO reactant on S-MSW pyrolysis are experimentally investigated. The content mainly consists of three parts: effect of temperature and in-furnace CaO reactant 1) on pyrolytic product properties; 2) on HCl mitigation efficiency; and 3) on tar cracking and tar composition upgrading.

Accordingly, the structure of this chapter is presented as follows:

I. Section 6.2 is dedicated to evaluate the effect of temperature and in-furnace CaO reactant on pyrolytic products mass distribution, produced gas yield and composition, and H 2 gas yield; IV. Section 6.5 summarizes the findings of this chapter.

Thermochemical conversion characteristics of S-MSW pyrolysis

The influences of temperature and in-furnace CaO reactant on product mass distribution, total gas yield, produced gas composition, and H 2 yield are discussed in this section. reactions between CaO and pyrolysis products, such as CO 2 , HCl, and etc., to form solid phase products (CaCO 3 , CaCl 2 , CaClOH, and etc.) may be responsible for the gradually increase of the solid residues yield. 6.2 summarizes some important thermochemical reactions reflecting waste pyrolysis and the role of CaO reactant [15,16,23,24].

Pyrolytic product (produced gas, tar, and solid residues) mass distribution

Produced gas yield and combustible gas composition

As exhibited, with the increase of temperature from 550 to 850 °C, H 2 molecular fraction in the produced gas is continuously increased from 8.3 to 18.3 %. According to [25],

at lower pyrolysis temperature, H 2 is mainly produced from decomposition of volatiles and the cracking reactions of aliphatic however, at higher temperature, the direct generation is mainly coming from dehydrogenation condensation-polymerization reactions of aromatic Chapter 6: Influence of CaO reactant on pyrolysis of simulated MSW: HCl mitigation and tar upgrading 159 hydrocarbons. Therefore, the formation of H 2 gas is significantly enhanced at higher temperature. Moreover, reforming reactions also play important roles on H 2 yield during waste pyrolysis. The enhancement of endothermic reactions at higher temperature, such as the general pyrolysis reaction (as presented in Table 6. 2, Eq. (6.1)) and the water-gas reactions (Eq. (6.2) and Eq. ( 6.3)) also facilitate the generation of H 2 gas. With regard to the addition of CaO reactant, H 2 yield is promoted by 0.8 -5.5 % under the experimental temperature. It is mainly because CaO could promote both tar catalytic thermal cracking reactions (Eq. (6.12)) and condensation-polymerization reactions of aromatic compounds.

Similarly, CH 4 yield is also increased from 13.7 to 31.8 % when pyrolysis temperature is increased from 550 to 850 °C. The formation of CH 4 during pyrolysis is mainly due to two pathways: one is from the decomposition of volatiles and the other is the further cracking of the produced methyl, oxymethylene, polymethylene and aliphatic hydrocarbons components [25]. Increasing the temperature apparently increases the generation of CH 4 . The use of CaO firstly enhance the CH 4 yield, mainly due to the enhancement of tar catalytic thermal cracking reaction (Eq. (6.10)); but at higher temperature, CH 4 content declined slightly because of the strengthening of hydrocarbon reforming reactions by CaO reactant.

The content of CO is varying in the range of 16.7% and 23.6% with temperature. The reason may be attributed to the competing reactions between water-gas reactions (Eq. (6.2)

and Eq. (6.3)) and water-gas shift reaction (Eq. (6.4)). With the addition of CaO reactant, CO content is limited. That can be explained by the enhancement of water gas shift reaction (Eq. (6.4)) due to the absorption of CO 2 by CaO. CO 2 in the produced gas is consumed through CaO carbonation reaction (Eq. (6.7)), leading to the acceleration of the water-gas shift reaction towards the direction of producing more H 2 and consuming CO.

Regarding C 2 Hx (the sum of C 2 H 2 , C 2 H 4 and C 2 H 6 ), the content increased dramatically from 8.8 to 24.7 % when increasing the temperature from 550 to 650 °C, but then declined slightly to 19.1 % when further increasing the temperature. The initial increase of C 2 Hx content is mainly attributed to the enhancement of tar thermal cracking (Eq. (6.6)) at higher temperature; however, further increasing the temperature favors the hydrocarbon reforming reactions (Eq. (6.5)), thus resulting in the decline of C 2 Hx content at higher temperature.

Similar to CH 4 , the yield of hydrocarbons (C 2 H x ) is firstly enhanced with the addition of CaO, due to the enhancement of tar catalytic thermal cracking reaction; but then declined slightly because of catalytic hydrocarbon reforming reactions.

Chapter 6: Influence of CaO reactant on pyrolysis of simulated MSW: HCl mitigation and tar upgrading H 2 gas yield is a key indicator to reflect the quality of the produced gas for further utilization. Therefore, the variation of H 2 yield as a function of temperature and CaO reactant is presented in Figure 6.2. As shown, the generation of H 2 is dramatically enhanced from 6.76

H 2 yield from S-MSW pyrolysis without/with CaO reactant

Chapter 6: Influence of CaO reactant on pyrolysis of simulated MSW: HCl mitigation and tar upgrading 161 to 79.52 mL/g-MSW when increasing the temperature; and the presence of CaO reactant increase the H 2 yield by approximately 25% to 88% under the experimental temperature.

Summary of the section

This section focuses on the role of CaO reactant and the effect of temperature on MSW pyrolysis products mass distribution, gas yield and composition, and H 2 yield. Results reveal that the addition of CaO reactant and the increase of temperature enhance the generation of gas products but decline the yield of tar compounds. In regard to combustible gas composition, H 2 yield is dramatically increased from 6.76 to 79.52 mL/g-MSW when increasing the temperature from 550 to 850 o C; and the presence of CaO reactant increases the H 2 yield by approximately 25% to 88% under the experimental temperature. With respect to CO, the content is fluctuant with temperature increase and CO generation is limited when CaO reactant is used. occurred to form CaCO 3 and CaClOH under the experimental conditions. Similar results are obtained by Yingjie Li et al [26], stating that CaCO 3 and CaClOH are the main products of CO 2 /HCl adsorption by Ca-based sorbent. According to Xin Xie et al [24], CaClOH is the main chlorination product formed with CaO reactant for short reaction time (at around tens of minutes), while CaCl 2 is the final chlorination product for long reaction time (at around hundreds of minutes).

Effect of temperature and in-furnace

As depicted in Figure 6.3, the intensity of CaO diffraction peaks increased significantly when increasing the temperature from 650 o C to 850 o C, while the intensity of CaCO 3 peaks declined simultaneously. It is mainly attributed to the fact that higher temperature shifts the thermodynamic equilibriums of carbonation/calcination reaction (Eq. (6.7)) towards the backward direction to decompose the crystalline CaCO 3 and to form more crystalline CaO.

Meanwhile, the intensity of CaClOH diffraction peaks also decreased when the pyrolysis temperature reaches 850 o C, indicating a reduction of HCl mitigation capacity at higher temperature to form crystalline CaClOH.

Effect of temperature and CaO on HCl mitigation efficiency

S-MSW pyrolysis

As aforementioned in Section 6.3.1, the addition of in-furnace CaO reactant captures the produced HCl gas to form CaClOH, thus immobilizing the chlorine into the solid residues, and prevents its release to the produced gas. With respect to the effect of Ca/Cl molar ratio, the stoichiometric Ca/Cl molar ratio of the HCl mitigation reactions is 1:2 and 1:1 respectively according to (Eq. (6.8) and Eq. (6.10))

and (Eq. (6.9) and Eq. (6.11)). However, HCl mitigation efficiencies under such molar ratios are quite low and only in the ranges of 42.2 -69.6% under 650 and 750 o C. This indicates that the excessive use of CaO reactant than theoretical stoichiometric amount is essential to achieve acceptable removal efficiency levels. The removal efficiency raise rapidly by approximately 40% when increasing Ca/Cl ratio from 1:2 to 2:1; however, further increasing the ratio to 4:1 lead to a limited enhancement by around 5%. Therefore, the Ca/Cl molar ratio of 2:1 is believed the most effective amount of CaO for HCl gas removal under the experimental conditions. Results reveal that both increasing the temperature and the addition of CaO significantly reduce the yield of pyrolytic tar content. As aforementioned in Chapter 5, tar decomposition is accelerated through both tar thermal cracking reactions (Eq. (6.6)) at higher temperature and tar catalytic thermal cracking reactions (Eq. (6.12)) with the addition of CaO reactant. in a limited enhancement of tar removal efficiency by approximately 2.5%. It is mainly attributed to the fact that at higher temperature as 850 o C, a great deal of tar is thermal cracked and the role of CaO reactant on tar cracking is not as important as that at lower temperature.

Morphology analysis of the spent in-furnace CaO reactants

As depicted in

At temperature of 750 o C, pyrolytic tar yield reaches 0.31 and 0.15 g/g-S-MSW without and with the addition of CaO reactant; tar thermal reduction efficiency is 34.6% and tar catalytic thermal reduction efficiency achieves 67.5%.

To summarize, the most preferable working condition concerning both tar catalytic thermal cracking and HCl mitigation efficiency as discussed in Section 6.3 is at temperature of 750 o C and Ca/Cl molar ratio of 2:1. Under such circumstance, HCl mitigation efficiency is 83.6%, pyrolytic tar yield is 0.15 g/g-S-MSW, and tar catalytic reduction efficiency reaches 67.5%. 6.3 and Table 6.4. As aforementioned in Section 5.2.1, the relative content of pyrolytic tar compounds are normalized and presented as Peak area% (the calculation method can be found as Eq. (5.1)). However, at higher temperature of 750 o C, no aliphatic hydrocarbons are observed in the detected tar compounds and only aromatic compounds are identified. The reasons may be two folds: on the one hand, higher temperature facilitates the decomposition and reforming reactions of aliphatic hydrocarbons to form light gases such as CH 4 , H 2 , CO, and CO 2 ; on the other hand, dehydrocyclization reactions are accelerated at higher temperature to form aromatic compounds. Therefore, under the experimental conditions, aliphatic hydrocarbons are consumed and only aromatic compounds are found in pyrolytic tar at temperature of As illustrated, at 550 o C, the main compounds are C11-C20 both in the absence of CaO reactant and with the addition of CaO, accounting for more than 79 % and 84% of the total relative content, respectively. The addition of CaO reactant slightly decreases the relative content of heavy compounds: no >C20 is identified with the addition of CaO, while without CaO, the content is approximately 4%; C16-C20 is reduced from 43% to 37% with CaO reactant. With regard to the relative distribution of organic categories, the addition of CaO reactant increases the content of both hydrocarbons (from 52% to 58%) and aromatic compounds (from 3% to 17%), but significantly decreases the relative content of oxygenated compounds from 39% to 25%, indicating that CaO reactant is effective to upgrade the quality of pyrolytic tar by deoxygenation reactions and thus reduce the oxygen content in tar compounds. Similar observations are reported by Zhu et al. [27], stating that tar produced are deoxygenated and desulfurized with the presence of CaO during coal pyrolysis.

As depicted in

In regard to the experiment conducted at temperature of 750 o C, only aromatic compounds are detected by GC/MS. It is observed that the carbon atom numbers of the major components are C11 -C15 under both non-catalytic and catalytic conditions. The compounds are a little bit lighter than that identified at 550 o C, where the major compounds are C11 -C20. With the increase of temperature from 550 to 750 o C, C16 -C20 species reduced dramatically from 43 -11% without CaO addition and 37 -7% in the presence of CaO reactant. This phenomenon directly demonstrates that thermal cracking and dehydrocyclization reactions of tar compounds occur with the increase of pyrolytic temperature. Similarly, the addition of CaO reactant slightly reduces the relative content of heavy compounds. C16 -C20 is reduced from 11% to 7% and C11 -C15 is declined from 54% to 52%. With respect to specific species, naphthalene and substituted naphthalene compounds are increased due to the addition of CaO reactant while the relative content of benzene is slightly decreased.

Summary of the chapter

Pyrolysis of S-MSW is experimentally investigated in this chapter. The effect of temperature and CaO reactant on pyrolytic product mass distribution, gas product characteristics, solid products characteristics, HCl gas removal and tar cracking and upgrading are determined. The main findings are as follows:

 Temperature is the determining factor affecting product yields by shifting the thermodynamic equilibriums of related chemical reactions. Higher temperature favors the generation of produced gas but declines the generation of tar and solid residues. When increasing the temperature from 550 to 850 o C, gas yield increased sharply from 81.9 to 451.4 mL/g-S-MSW. The presence of CaO increased the total gas yield by 30.0 to 58.8 mL/g-S-MSW and enhanced the H 2 yield by approximately 25% to 88% under experimental temperature.

 HCl is captured by CaO reactant and retained in solid phase products.

Immobilization of chlorine can thus be achieved by using CaO reactant to avoid the presence of HCl in the produced gas. CaClOH is identified as the main dechlorination product by XRD observations. Increasing the temperature causes adverse effects on HCl mitigation efficiency and the increasing of Ca/Cl molar ratio promotes the efficiency.

 Both increasing the temperature and the use of in-furnace CaO reactant are effective ways to reduce pyrolytic tar content. At lower temperature, aliphatic hydrocarbons are the major compounds of pyrolytic tar, while at higher temperature, only aromatic species are detected. Increasing the temperature decomposes the heavy compounds and facilitates the formation of aromatic by thermal cracking and dehydrocyclization reactions. The use of in-furnace CaO reactant reduces the oxygen content in tar compounds and improves the content of naphthalene and substituted naphthalene compounds.

 The recommended working condition concerning both tar cracking and HCl mitigation efficiency from the present work is: temperature of 750 o C and Ca/Cl molar ratio of 2:1. Under such circumstance, HCl mitigation efficiency is 83.6%, pyrolytic tar yield is 0.15 g/g-S-MSW, and tar catalytic reduction efficiency reaches 67.5%.

Chapter 7

Life cycle assessment of MSW thermal technologies with different dehydrochlorination options

Introduction

Municipal solid waste (MSW) thermal technologies have become one of the most promising options to achieve both waste treatment and energy recovery [1]. MSW incineration, for example, is playing an increasingly important role and is developing rapidly worldwide in the past few decades [2][3][4][5]. By incineration, mass and volume of MSW can be dramatically reduced and its heating value can be recovered to generate electricity and/or heat [6][7][8]. However, the release of harmful pollutants, especially NO x , SO x , dioxins and heavy metals from stack causes severe environmental burdens and seriously threatens human health [9]. In addition, as previously mentioned, chlorine in MSW causes serious problems during MSW incineration, related to corrosion, toxic organics formation, and acidification, etc.

Instead of converting all the energy contained in MSW to thermal energy in flue gas by incineration, pyro-gasification technology allows the generation of fuel gases (syngas) and bio-oil (tar) from MSW that can be further utilized to create heat and/or electricity or alternatively as feedstocks to produce fuels or chemicals [10][11][12].

As stated by Chapter 1, thorough and systematic evaluation of current MSW thermal technologies is of great importance to determine a more energy efficient and environmental friendly waste to energy (WtE) route. As a result, this chapter is focused on the assessment and optimization of current MSW thermal conversion technologies using LCA methodology.

In addition, special attention has been paid to pyro-gasification-based WtE systems with different dehydrochlorination strategies (for example ex-situ low temperature dehydrochlorination; ex-situ high temperature dehydrochlorination, and in-situ dehydrochlorination). The aim is to provide scientific basis for decision makers regarding the improvement of local waste management plan. With this overall objective, two LCA studies are performed: -Section 7.2 is mainly dedicated to give a systematic and holistic comparison between current commercial MSW thermal conversion technologies including conventional incineration, pyrolysis, gasification-melting, and gasification with ex-situ high temperature dehydrochlorination as performed by Lahti Energia; environmental burdens caused by different systems as well as the main contributor process are identified in this section; -Section 7.3 is mainly focused on the environmental performance of pyro-gasification-based WtE systems with different dehydrochlorination strategies, including ex-situ low temperature dehydrochlorination (gasification-melting); ex-situ high temperature dehydrochlorination (gasification, Lahti), and hypothetical in-situ dehydrochlorination systems; the aim is to determine the role of dehydrochlorination strategy in WtE systems; -Section 7.4 conducts the sensitivity analysis; key parameters adopted in this study are varied to examine the effect of data use on the final results; -Section 7.5 summarizes the chapter.

LCA of four current commercial technologies

Guided by ISO standards 14040 series [13], the general framework of LCA studies consists of four phases: (1) goal and scope definition; (2) life cycle inventory (LCI); (3) life cycle impact assessment (LCIA); and (4) interpretation. These four steps are described separately below.

Goal and scope definition

According to literature [10,14], there are more than 100 gasification-based WtE plants around the world so far; and most of them are located in Japan, generally based on gasification-melting technology; or on a small scale, based on pyrolysis. In Europe, the implementation of pyro-gasification-based WtE technologies have become more and more popular in recent years, such as Lahti in Finland [15], Averoy in Norway [16], and waste pyrolysis plant Burgau in Germany [17]. Table 7.1 generally summarizes some of the waste gasification plants running in Europe [18]. Greve-in-Chianti, Italy pelletized RDF Gasification Norrsundet, Sweden MSW and industrial waste Gasification

Among those applications, it is observed that most of the pyro-gasification processes are linked to a downstream energy recovery device, or k o as t o-step o idatio WtE scheme [10,[19][20][21]. One of the advantages of pyro-gasification technologies is the possibility to achieved energy efficient and environmental friendly utilization of the produced gas rather than directly incinerate the heterogeneous waste.

Therefore, the primary aim of this section is to comprehensively evaluate the environmental performance of pyro-gasification-based WtE systems and to compare them with typical conventional MSW incineration scenario. As a result, a total of four systems are considered in this section, with the LCA system boundaries illustrated in transportation before entering the system boundary are excluded from this, as they are identical in all systems [22]. Wastewater effluent during treatment is not considered due to the lack of data; however this assumption will not cause significant deviations since modern incinerators are designed with wastewater treatment, recycling and reused equipment in the pla t to eet the ze o dis ha ge ta get [23,24]. Besides, energy/emission due to plant construction and demolition is also omitted, as it is negligible compared to the total burdens aroused by plant operation [25].

Not only the direct energy/material consumption and emission are considered, the upstream and downstream input and output from background system are also taken into account in this study. Acquisition of auxiliary materials and energy, such as diesel, lime, etc., is also o side ed; thus, the al ulatio t pe of adle to g a e is a hie ed. The output electricity from the systems are used to substitute the same amount of electricity from the E.U. average electricity grid, and thus the corresponding environmental emissions to produce the same amount of electricity is considered to be avoided.

One tonne of MSW (as received basis) is defined as the functional unit. Therefore, inventory data of all input and output materials, energy and emissions are converted to this basis. MSW characteristics are determined in accordance with typical values of the E.U.

average MSW fractions, and are shown in Table 7.2 [26,27]. Background data related to raw material production are mainly obtained from Gabi software. denitrification with use of reagents for dioxins abatement, dry process for acid gas removal, and bag filter for particles removal [2,8]. Metal content in the bottom ash is recovered and the rest of the residues (APC residues are firstly solidified/stabilized) are disposed in landfill site.

S2: MSW pyrolysis system

S2 refers to the MSW pyrolysis system. The selected plant Bu gau MW P ol sis Pla t , located in Günzburg, Germany, has a capacity of 20,000 t/a [29]. After shredded to 30 cm, the MSW is decomposed in the absence of air in an indirect heated rotary kiln at temperature of 500-600 °C with a residence time of 1 h. The rotary kiln is heated indirectly with flue gas from the pyrolysis gas incineration. The pre-dedusted pyrolysis gas is extracted and charged to a combustion chamber to be incinerated with an excess air of approx. As aforementioned in Chapter 1, S3 represents gasification system with ex-situ high temperature purification of the produced syngas as performed by Lahti Energia [30,[START_REF] Coda | Behavior of chlorine and enrichment of risky elements in bubbling fluidized bed combustion of biomass and waste assisted by additives[END_REF]. The plant is based on circulating fluidized bed (CFB) for the production of electricity (50 MW) and district heat (90 MW). Its commercial operation was started in 2012, and approximately 250,000 tons of SRF can be handled annually [32]. The gasifier runs at 850-900 o C under atmospheric pressure. As a result of gasification, the feedstock (SRF) turns into produced syngas. The produced syngas then undergoes a cooling and purification step at 400 o C. This temperature is chosen because on one hand, corrosion caused impurities (e.g. alkali chlorides) can condense as solid particles; on the other hand, avoid tar condensation. Heat from the gas cooling step is recovered to preheat the feed water. The cleaned syngas passes through a filter to remove particles, before combustion in a secondary chamber at 850 o C.

The flue gas is relatively clean to bear a higher quality of steam superheated in the boiler (540 o C, 121 bar). The electricity generation efficiency is calculated as 27.4% based on the technical report. And the same APC systems are assumed as for the incineration case. For solid residues disposal, the produced bottom ash is landfilled; while the fly ash and air pollution control (APC) residues require safety stabilization before final disposal

S4: MSW gasification-melting system

As previously mentioned, MSW gasification-melting plants is the most popular gasification-based WtE technologies. The main reason to choose this kind of technology is its possibility to effectively recover materials [33]. The selected plant, having a total throughput of 80 MW, is located in Japan and is one of the largest gasification-melting facilities in the pre-treatment of the incoming waste is required. The syngas is combusted to generate steam at 400 °C and 3.92 MPa. The net electricity efficiency reaches 17% [33]. The APC system is considered the same as that adopted in incineration system. The molten materials from the gasifier are magnetically separated into slag and metals, which can be completely recycled.

Pretreatment of MSW

Since SRF is used in gasification scenario (i.e. S3), energy and materials consumption during SRF production is also considered as input streams. According to previous works [34,35], energy consumption in the SRF production process mainly includes electricity and fuels and is estimated at 804 MJ of electricity and 32.3 MJ of diesel per ton of MSW.

Storage for dehydration and shredding for size reduction of MSW is also important for other systems; however, electricity consumption in this stage is considered as internal electricity consumption of the power plant and has been considered in this study.

Energy and materials recovery

In the present work, only electricity generation is considered as energy recovery, which means heat recovery or combined electricity and heat (CHP) system is not taken into account.

For materials recovery, ferrous metals and aluminum recovery from bottom ash is included in this work.

Life cycle inventory

The compiled life cycle inventory (LCI) of the present LCA study is displayed in Table 7.3.

Besides, some key assumptions are made to fulfill the calculation and are listed as follows:

 Only fossil-derived CO 2 emission is considered as a contributor to GW, biogenic CO 2 emission is regarded to be carbon neutral.

 The raw MSW compositions are considered identical for all the systems.

 Emissions during the ash disposal process are obtained from literature, in which statistical data that are currently used for dealing with MSW in the UK and in Europe are applied as the calculation basis [36].

Chapter 7: Life cycle assessment of MSW thermal technologies with different dehydrochlorination options As aforementioned, electricity generation is served to substitute the same amount of electricity from conventional electricity matrix (in this study, the E.U. average electricity matrix is adopted as the reference and it is presented in Table 7.4 [START_REF] Slapak | Determination of the pyrolytic degradation kinetics of virgin-PVC and PVC-waste by analytical and computational methods[END_REF]).

Life cycle impact assessment

As pointed out in Chapter 2, the CML (baseline) methodology is used to aggregate the LCI results [38]. Six impact categories are considered, including three non-toxic and three environmental savings) can always be obtained for the environmental impact categories of AP, FAETP, and HTP from the four considered systems; however, positive impacts (i.e. In order to identify the main sources of environmental burdens and benefits, the overall environmental impacts are distributed into eight life cycle stages, including: stack emissions, materials and energy provision, bottom ash landfill, slag recycling, APC residues stabilization and landfill, metals recycling, residues transportation, and energy recovery. shows also a further avoided contribution, in particular for toxicity impacts regarding FAETP and HTP.

With respect to specific impact categories, the environmental burdens of GWP is mainly aused sta k e issio , accounting for approximately 417, 420, 420, and 602 kg CO 2 eq. of " , " , " , a d " s ste , espe ti el . It is o th oti g that GWP f o sta k e issio of S1, S2 and S3 systems is quite similar (ranges from 417 -420 kg CO 2 eq.), but the value is much higher from S4 system. The reason is mainly attributed to the use of coke as auxiliary fuel in S4 system, leading to much higher fossil CO 2 generation and more contribution to GWP [33]. thus not recyclable in S4 [39].

For the toxic impacts of FAETP and HTP, sta k e issio is no longer the key o t i uto of the e i o e tal u de s; i stead etals e li g sho s significantly avoided contribution on FAETP and HTP. For the impact of FAETP, the recycling of ferrous metals accounts for more than 96% of the total credits. The benefit comes from the saving of virgin metals production, which indicates, the environmental impacts for scrap metals separation, upgrading, and reprocessing are considerably lower than their production from virgin materials. S4 leads to the best FAETP savings, mainly attributed to the highest ferrous metals recovery of 12.60 kg as presented Table 7.3. For the impact of HTP, the credits are nearly completely ascribed to Al recycling. Therefore, as aforementioned, HTP savings from S4 system is the worst among the four systems and S2 performs the best due to the highest Al recovery of 6.82 kg, followed by S1 (Al recovery of 6.26 kg) and S3 (Al recovery of 4.64 kg).

Summary of the section

In this section, environmental impact from four considered MSW thermal conversion systems are evaluated by LCA methodology. The considered systems include the conventional MSW incineration (S1, the baseline case) and pyro-gasification based WtE systems (S2, pyrolysis-based system; S3, gasification-melting based system; and S4, gasification based system). Results reveal that the environmental performance of the considered systems in descending order is: S3, Gasification > S2, Pyrolysis > S4, Gasification-melting > S1, Incineration. In regard to the contributor of the environmental impacts, sta k e issio s are the most important environmental burdens for the impact of GWP, AP, EP, and TETP. E e g e o e makes a significant contribution to reduce the GWP, AP, and EP impacts. And metals e li g sho s also a fu the a oided o t i utio , i particular for toxicity impacts regarding FAETP and HTP.

Gasification-based WtE system with different dehydro-chlorination strategies

In this section, a simplified, idealized, and hypothetical gasification-based WtE system using in-furnace dehydrochlorination strategy is proposed on basis of S3 system. Assumption has been made that:

Chapter 7: Life cycle assessment of MSW thermal technologies with different dehydrochlorination options

193

-By adding in-furnace CaO reactant, all the released chlorinated contaminants are absorbed and stored into the solid residues, which means, no chlorine-content product (HCl, Cl 2 , chlorides, chlorinated organics, etc.) contained in the produced gas. Therefore, the produced gas cooling down and purification stage (as applied in S3) of the hypothetical system is not essential. And the steam parameters can achieve as high as that used in S3;

-The use of in-furnace CaO acts as CO 2 absorption sorbent to decrease the fossil CO 2 emission; however, the catalytic effect on produced syngas quality enhancement is not taken into account due to the lack of available data;

-Due to the assumption of completely Cl abatement, no HCl and dioxins are discharged at the stack.

The aim of this section is to estimate the potential environmental benefits from complete in-situ dehydrochlorination during MSW gasification. The systems considered and the corresponding system boundaries in this section are presented in a the consumption of electricity and diesel is increased because more bottom ash is produced due to the addition of in-furnace CaO;

b according to Jhong-Lin Wu et al [40], chlorine content in MSW ranges from 800 -2500 mg/kg. The value of high chlorine content as 2500 mg/kg is adopted in this study. The in-furnace CaO amount is calculated based on Ca/Cl molar ratio of 2:1 according to the conclusions obtained from Chapter 4 and Chapter 6;

c due to the in-furnace removal of chlorinated contaminants, produced gas cooling down and purification stage is not essential in the proposed system. Therefore, the apparent energy of the produced gas can be utilized to generate electricity. The calculation is based on our previous work [21], revealing that hot gas efficiency is approximately 15% higher than cold gas efficiency during MSW gasification at 850 o C; Taking into account the electricity efficiency of S3 (27%), the increase of the electricity generation as compared to S3 is approximately 4%; e due to the assumption of complete Cl removal, no HCl and dioxins is considered;

f CO 2 emission is slightly decreased due to CO 2 absorption by CaO;

g the rest of the parameters are adopted from S3.

Accordingly, the characterized environmental impacts of the considered system are listed in Table 7. 7 and the contribution analysis results are illustrated in Figure 7.9. Results reveal that as compared to S3 (Gasification system) and S4 (Gasification-melting system), the proposed hypothetical gasification-ased s ste " leads to the est environmental performance on the impact categories of GWP, AP, TETP, and HTP; while the S4 system still causes excellence behavior over the impact of EP and FAETP (as aforementioned in Section 7.2). It is o th e tio i g that GWP f o " is sig ifi a tl reduced compared with S3; the reason may be attributed to the increase of electricity generation from the hypothetical system. Actually, it is observed that when compared with S3, all the environmental u de s f o " a e de li ed, which means, from the present study, the proposed hypothetical gasification-based system using in-situ dehydrochlorination strategy has a better environmental performance than gasification system with ex-situ dehydrochlorination. However, the obtained results are mainly based on the assumptions that most of the ope atio o ditio s a d pa a ete s of " a e si ila to that of " ; ut approximately 4% of the electricity is additionally generated owning to energy savings from produced gas apparent energy. Further consideration and examination of the parameters adopted and the process system should be investigated on these two complex systems.

Summary of the chapter

In this chapter, LCA investigations of four WtE thermal conversion systems (including S1, conventional waste incineration system; S2, pyrolysis-based WtE system; S3, gasification system with ex-situ dehydrochlorination as performed by Lahti Energia; and S4, gasification-melting technology) are firstly conducted to determine the environmental performance of pyro-gasification-based WtE systems when compared with conventional MSW incineration. Then a simplified LCA comparison between gasification-based systems with different dehydrochlorination strategies is carried out. The main findings are as follows:

 Gasification system (S3) leads the least environmental impacts for most of the considered categories such as GWP, AP, and TETP; but it also causes the highest environmental burdens of FAETP. Gasification-melting system (S4) performs worst in the environmental categories of GWP, TETP, and HTP; but S4 system is efficient to reduce the environmental impact of EP and FAETP;

 The environmental performance of the considered systems in descending order is: S3, Gasification > S2, Pyrolysis > S4, Gasification-elti g > " , I i e atio > means the former system performs better than the latter one).

 In regard to the contributors of the overall systems, stack emissions are the most important environmental burdens for most of the impact categories; energy recovery and metal recycling account for most of the environmental savings.

Mate ials a d e e g p o isio akes o l a i o o t i utio to " a d " scenarios, however it is more noteworthy in S3 and S4, for pretreatment of MSW and the addition of auxiliary fuel, respectively.

 The proposed hypothetical gasification-based system (S3 ) leads to the best environmental performance on the impact categories of GWP, AP, TETP, and HTP; while gasification-melting based system (S4) still causes excellence behavior over the impact of EP and FAETP. It is concluded that the environmental performance of S3 is superior to that of S3. However, it should be mentioned that the evaluation is mainly based on assumptions and it is an idealized and simplified estimation.

Therefore, the parameters adopted and the whole system flow sheet of the hypothetical gasification-based system with in-situ dehydrochlorination should be further identified and examined.

Results reveal that increasing the temperature, reducing raw concentration of HCl gas, and the presence of CO, and H 2 in carrier gas causes the reduction of HCl mitigation capacities and CaO bed conversion efficiency. Higher temperature shifts the chemical equilibrium of the absorption reaction and reduces surface area and pore volumes of CaO reactant, thus resulting in a decreasing of HCl mitigation capacity from 778.9 to 173.9 Results reveal that the gasification system (S3) leads to the least environmental impacts for most of the considered categories such as GWP, AP, and TETP; but it also causes the highest environmental burdens of FAETP. Gasification-melting system (S4) performs worst in the environmental categories of GWP, TETP, and HTP; but S4 system is effective to reduce the environmental impact of EP and FAETP. The environmental performance of the considered systems in descending order is: S3, Gasification > S2, Pyrolysis > S4, Gasification-melting > S1,

I i e atio > ea s the fo e s ste pe fo s better than the latter one).

Then a simplified LCA comparison between gasification-based systems with different dehydrochlorination strategies (S3, gasification system using ex-situ produced gas purification; S4, gasification-melting system using APC systems for acid gas removal; and S3 , the hypothetical gasification-based system using in-situ dehydrochlorination) is carried out.

Results reveal that the proposed hypothetical gasification-ased s ste " leads to the best environmental performance on the impact categories of GWP, AP, TETP, and HTP; while gasification-melting based system (S4) still causes excellence behavior over the impact of EP a d FAETP. It is o luded that the e i o e tal pe fo a e of " is supe io to that of " .

However, it should be mentioned that the present work is mainly based on assumption and is an idealized and simplified estimation. Therefore, the parameters adopted and the whole system flow sheet of the proposed hypothetical gasification-based system should be further identified and examined.

Prospects

Pyro-gasification-based MSW thermal conversion system is attracting more and more concerns due to the potential to achieve energy efficient and environmental sound waste to energy routes. However, there are several bottlenecks needs to be overcome for its popularization and worldwide applications. Chlorine content is one of the biggest challenges during MSW thermal conversion. Besides, new challenges will also come up from both technological and theoretical evaluation perspectives. This section is dedicated to the prospects for future research investigations.

 Improvement of reactant properties for high temperature HCl mitigation

Results from this study indicated that the use of in-furnace CaO reactant is effective to absorb the released HCl during MSW pyro-gasification; however, the removal capacity declined with the increase of temperature. Therefore, in-depth investigation in this field is essential to fulfill the research, which could include:

(i) Enhancement of reactant properties. In this study, only the base reactant CaO is investigated for both HCl mitigation and pyrolytic tar upgrading. It is concluded that the excessive use of CaO than the stoichiometric amount is essential to achieve acceptable HCl mitigation efficiency. Therefore, further investigations could be focused on the enhancement of reactant properties, such as pretreatment of CaO [1], reactive metals loading on CaO surface [2,3], and mixing with other reactants [4];

(ii) Re-generation or use of Ca-content waste as in-situ reactant. Re-generation of spent reactant is of great importance to achieve environmental sustainability [5]; but it is not investigated in the present work. Therefore, further study work could be focused on this aspect. Another option can be focused on the reuse of Ca-content waste, for example carbide slag, as indicated by previous studies [6,7];

(iii) Development of high temperature HCl mitigation reactant. As aforementioned, the capacities of CaO reactant on HCl mitigation declines rapidly with the increase of temperature. Therefore, the development of new and effective reactant, which works at higher temperature windows, is also a fascinating research field [8].

Chapter 9

French summary of the thesis La confrontation des résultats expérimentaux avec les modèles cinétiques montre que l ́tape i ́ti ue d́te i a te est la ́a tio hi i ue h́t́ ogè e à la su fa e du CaO o réagi), puis la diffusion de HCl à travers la couche de produit formée (CaCl 2 ). Entre ces 2 ́gi es, u e zo e de ́gi e i te appa aît do t l ́te due d́pe d de la te ṕ atu e.

Figure 1 . 1

 11 Figure 1. 1 MSW generation in France and China during 2007-2010,data derived from[1]

Figure 1 . 3

 13 Figure 1. 3 Share of MSW treatment methods in the EU during 2004-2015, data source from Eurostat database[7]

Figure 1 . 5

 15 Figure 1. 5 The main ideas of this study: investigate in-furnace capture of chlorine to avoid the release of HCl by using in-situ CaO reactant

  easily and thoroughly during thermal conversion. Pyrolytic degradation of PVC experiments revealed a two-stage reaction mechanism associated with: (1) thorough dehydrochlorination (de-HCl) occurring in low temperature range of 200 o C -400 o C; and (2) further degradation of the de-HCl PVC polyene chains taking place at temperature range of 400 o C -550 o C

Figure 1 . 6

 16 Figure 1. 6 Graphic principle of acid rain formation pathway: (1) Emissions of SO 2 , NO x and HCl are released into the air, where (2) the pollutants are transformed into acid particles that may be transported in a long distance, (3) the acid particles then fell to the earth as wet and dry deposition (dust, rain, snow, or etc.) and (4) may cause harmful effects on soil,

Figure 1 . 7

 17 Figure 1. 7 Schematic diagram of high temperature corrosion occurred on the surface of the equipment, data obtained from Wang[START_REF] Ting | Study of calcium oxide dechlorination in biomass fluidized bed gasifier[END_REF]

Figure 1 . 8 Figure 1 . 8

 1818 Figure 1. 8 Spray Drying Gas Cleaning System, data obtained from Razvan Lisnic et al. [74]

Figure 1

 1 Figure 1.9.

Figure 1 . 9

 19 Figure 1. 9 Phases and applications of an LCA standardized by ISO, data obtained from the work of Jun Dong[START_REF] Dong | MSWs gasification with emphasis on energy, environment and life cycle assessment[END_REF]

Figure 1 .

 1 Figure 1. 10 LCA concerns in waste management field, data source from Laurent et al.[START_REF] Laurent | Review of LCA studies of solid waste management systems-Part I: Lessons learned and perspectives[END_REF]

Figure 1 .

 1 Figure 1. 11 Structure and main contents of the thesis

Section 2 . 1 -Section 2 . 3 Chapter 2 :Figure 2 . 1

 2123221 Figure 2. 1 Pictures of MSW components: (a), PVC pellet; (b), wood; (c), food waste; (d), PE pellet; and (e), paper

Figure 2 . 1

 21 Figure 2.1 presents the pictures of MSW components. Before the experiment, feedstocks are ground into sizes of less than 1 mm and then dried at 105 o C for 24 hours. Physical composition and the ultimate and proximate analysis results of S-MSW components as well as the prepared S-MSW are presented in Table 2.1 and

Figure 2 . 2 Figure 2 . 2 . 3 .Figure 2 . 3 NChapter 2 :

 22223232 Figure 2. 2 XRD analysis of CaO reactant

Figure 2 . 4 Figure 2 . 4 .

 2424 Figure 2. 4 General photograph and schematic diagram of the vertical tubular furnace reactor: 1, gas source; 2, gas flow meter; 3, vertical quartz tubular reactor; 4, electric heater and temperature controller; 5, quartz filter; 6, CaO reactant for HCl mitigation; 7, quartz wool; 8, HCl gas detector

Figure 2 . 5

 25 Figure 2. 5 Schematic diagram of the experimental fixed bed tubular furnace apparatus. 1, gas source; 2, gas flowmeter; 3, quartz tubular furnace; 4, quartz boat; 5, electric heater and temperature controller; 6, condenser for tar collection; 7, cooling water; 8, HCl mitigation solution (0.1 mol/L NaOH); 9, silica gel; 10, ice-water bath

Figure 2 . 6

 26 Figure 2. 6 General photograph and schematic diagram of the Py-GC/MS experimental apparatus

Figure 2 . 5 ,

 25 0.1 mol/L NaOH is used to absorb the produced HCl together with Cl 2 (trace amount). Ice water bath is used to minimize the loss of absorbing solution, and at the end of the sampling train, silica gel is used to dry the produced gas and to protect the dry gas meter and pump. After the experiments, the absorbed solution is then quantitatively transferred to a volumetric flask and dilute with water. Ion chromatography (IC, Thermo Fisher Scientific) is used to determine the Cl -content of the diluted solution. The amount of HCl collected by the solution can be obtained by: HCl is the mole of HCl trapped by the absorption solution; C Cl -represents the concentration of Cl -detected by IC; V s is the volume of the diluted sample in the volumetric flask; M Cl-is the molecular weight of Cl -, 35.5 g/mol.

2. 4

 4 Thermodynamic simulation using FactSage software 2.4.1 General principle of thermodynamic equilibrium simulation

Figure 2 . 7

 27 Figure 2. 7 General principle of the thermodynamic equilibrium calculation system

Figure 2 . 8 .

 28 FactSage is believed one of the most commonly used chemical thermodynamics databases and also computing systems in the world[10, 11]. It is a fusion of two thermochemical packages named FACT-Win/F*A*C*T and ChemSage/SOLGASMIX by Thermfact/CRCT (Montreal, Canada) and GTT-Technologies (Aachen, Germany), respectively.

Figure 2 . 8

 28 Figure 2. 8 Main menu of FactSage 6.3 software

Figure 3 . 1

 31 Figure 3. 1 Schematic diagram of the reaction between CaO and HCl, showing the growth of the CaCl 2 product-layer and the consumption of unreacted CaO core

Figure 3 . 1 ,Figure 3 . 2 Figure 3 . 2

 313232 Figure 3. 2 Concentration of HCl under the condition that the reaction is only limited by gas film diffusion (A refers to HCl gas)

.Figure 3 . 3 Figure 3 . 3

 3333 Figure 3. 3 Concentration of HCl gas under the condition that the reaction is only limited by internal diffusion of HCl through the product-layer (A refers to HCl gas)

14 )Figure 3 . 4 Figure 3 . 4

 143434 Figure 3. 4 Concentration of HCl gas under the condition that the reaction is only limited by gas-solid chemical reaction (A refers to HCl gas)

15 )Figure 3 . 5 Figure 3 . 5

 153535 Figure 3. 5 Normalized reaction time as a function of CaO reactant conversion rate

Chapter 3 :

 3 Experimental and kinetics study on HCl mitigation by CaO reactant

  than 1980 ppm (similarly, 990 ppm for c 0 = 1000 ppm and 495 ppm for c 0 = 500 ppm is used to measure the saturation time);The total amount of HCl gas absorbed by one gram of CaO reactant is calculated according to: HCl/g-CaO represents the HCl mitigation amount by one gram of CaO reactant during the experiment; c 0 refers to the inlet concentration of HCl; c t is the instant outlet concentration of HCl at time t; V represents the gas flowrate; 36.5 is the molecular weight of HCl; V HCl is the molar volume of HCl; and m CaO is the mass weight of CaO reactant used. The HCl mitigation rate, which indicates the reaction rate between HCl and CaO during the mitigation process, can be estimated by taking into account the absorption amount of HCl over a period of time, and it is obtained by:

Chapter 3 : 73 3. 3 . 1 3 . 3 . 1 . 1

 373313311 Experimental and kinetics study on HCl mitigation by CaO reactant Effect of temperature on HCl mitigation by CaO reactant Breakthrough experimental and kinetics modeling resultsIn this section, the fixed parameters are: the raw HCl concentration is 2000 ppm, gas flowrate is 2.0 L/min and the amount of CaO used is 2.0 g. Under such conditions, the effect of reaction temperature on HCl mitigation is investigated. Experiments are carried out in moderate high temperature ranges from 550 to 850 o C. Figure 3.6 presents HCl breakthrough curves of CaO reactants under different temperature. Outlet HCl concentration is normalized by c/c 0, in which c represents the instant HCl concentration and c 0 refers to raw HCl concentration. Results reveal that increasing the temperature significantly decrease the breakthrough and saturation time during HCl mitigation using CaO reactant. When the temperature is increased from 550 to 850 o C, the breakthrough and saturation time decline from 197.8 to 21.8 min and 268.9 to 69.6 min, respectively.

  temperature will largely reduce HCl mitigation efficiency by CaO reactant from both physical adsorption and chemical reaction point of view. With regard to HCl mitigation capacity, Figure 3.7 exhibits the total HCl mitigation amount and the mitigation rate with time using CaO reactant within the considered reaction temperature ranges.

Figure 3 . 7 3 . 6 ,

 3736 Figure 3. 7 The total HCl mitigation amount and mitigation rate with time under different reaction temperature

Figure 3 . 8 Figure 3 . 9 .8 and 3 . 9 , 3 . 8 ,

 38393938 Figure 3. 8 Experimental and calculated P(x), C(x) vs CaO conversion rate curves at 550 o C

3. 3 . 1 . 2 Figure 3 .Figure 3 . 1 Figure 3 . 10 ,Figure 3 . 11 ,Figure 3 .Figure 3 .

 31233131031133 Figure 3. 10 Formation of the product-layer at temperature of 650 o C: a, t/t f = 0.1; b, t/t f = 0.3; c, t/t f = 0.5; d, t/t f = 1 (t f refers to the saturation time of the process)

Figure 3 . 4 .

 34 Figure 3.13 exhibits the micrograph of the final spent CaO reactant at different reaction temperature and their corresponding physical properties are listed in Table 3.4. As illustrated, the spent CaO reactant at reaction temperature of 550 o C has a higher specific surface area and total pore volume when compared with raw CaO. The reason may be attributed to the preparation of CaO reactant; as aforementioned in Section 2.2, raw CaO reactant is calcined at 800 o C for 2h to decompose impurities, and the surface area and pore volume are thus significantly reduced due to calcination. Therefore, raw CaO in the present work is considered nonporous and the shrinking unreacted core model is used to determine the reaction kinetics. HCl mitigation at 550 o C by CaO reactant generates a relatively high porous product-layer, thus increases the porosity of the sorbents. However, when the reaction temperature is increased from 550 to 750 o C, specific surface area and pore volume of the spent CaO reduces continuously from 4.2 to 1.5 m 2 /g and 4.4x10 -2 to 1.6x10 -2 cm 3 /g, respectively. The phenomena further confirm that the reduction of surface area and closure

Figure 3 .

 3 Figure 3. 14 Effect of raw HCl concentration on total HCl mitigation amount and HCl mitigation rate by CaO reactant

Figure 3 . 15 .

 315 Results reveal that the presence of CO 2 (gas containing 12% CO 2 ) slightly reduces the breakthrough time of HCl mitigation by CaO reactant from ca. 106 to 88 min. The competition of dechlorination reaction (as presented in Eq. (3.1)) and CaO carbonation reaction (Eq. (3.23)) is responsible for this phenomenon, since CO 2 may also occupy the active sites on CaO core surface and inhibit the reaction between HCl and CaO.However, the saturation time and total HCl mitigation amount is slightly promoted by approximately 10 min and 10 mg/g-CaO. One possible explanation may be attributed to the inverse calcination reaction of CaCO 3 (Eq. (3.24)), according to the literature[19], the decomposition of CaCO 3 can continuously produce active sites and increase pore volume and structures on the produced CaCO 3 layer, and thus enhance HCl mitigation.

Figure 3 .

 3 Figure 3. 15 Effect of gas composition on total HCl mitigation amount and HCl mitigation rate by CaO reactant

In chapter 3 ,

 3 Higher temperature shifts the chemical equilibrium of the absorption reaction and reduces surface area and pore volumes of CaO reactant; lower raw HCl concentration increases the breakthrough and saturation time, but decreases HCl mitigation amount and CaO bed conversion efficiency due to the reduction of both physical adsorption and chemical reaction rates; and the presence of CO, and H 2 will occupy the number of free active sites on CaO surface and cause adverse effects on HCl mitigation. The proposed reaction mechanisms are described as follows: Estimation results of kinetic modeling indicate that the ratio of product-layer diffusio esista e to he i al ea tio esista e i.e. θ 2 ) is always in the range of < θ 2 < 10, which means that HCl mitigation by CaO reactant, is limited by both chemical reaction and product-layer diffusion. Morphology analysis reveals the formation process of the product-layer and the examination of the spent CaO physical properties further indicates that both the shifts of chemical equilibrium and the reduction of surface area and closure of pore structures at higher temperature are responsible for the decline of HCl mitigation efficiency by CaO. absorption of HCl gas by CaO reactant and the gas-solid reaction mechanism are investigated. Results reveal that the use of CaO reactant is effective to capture HCl contained in the flue gas. In this chapter, the conversion of Cl-contained contaminants from inorganic and organic chlorine source in MSW is firstly identified and then the use of in-furnace CaO reactant on the produced HCl mitigation is examined.

Figure 4 . 1

 41 summarized the possible transformation routes of inorganic and organic chlorine sources during MSW thermal treatments of pyro-gasification and incineration.

Figure 4 . 1

 41 Figure 4. 1 Schematic diagram of the possible conversion routes of chlorine during MSW thermal treatments

Figure 4 . 2

 42 Figure 4. 2 Thermodynamic simulation results of pure NaCl under N 2 atmosphere

Figure 4 . 3

 43 Figure 4. 3 Thermodynamic simulation results of NaCl under considered working conditions: (a) under O 2 atmosphere; (b) under N 2 atmosphere with water; (c) under O 2 atmosphere with water; and (d) with the presence of SiO 2 and water under O 2 atmosphere

Figure 4 . 3

 43 Figure 4.3 (b). Results reveal that the presence of water promotes the evaporation of NaCl in

Figure 4 . 3

 43 Figure 4.3 (c) exhibits the effect of water on the thermodynamic behavior of chlorine from NaCl under O 2 atmosphere. Similar results are obtained as for water/N 2 atmosphere.

Figure 4 . 3

 43 (d). Results clearly demonstrate that the presence of SiO 2 significantly enhances the conversion of chlorine from NaCl to HCl.

Figure 4 . 4

 44 Figure 4. 4 TG analysis results of NaCl under N 2 atmosphere

Figure 4 . 4 ,

 44 Figure 4.4, NaCl is quite stable at lower temperature; no mass loss is observed before 700 o C.

Chapter 4 :

 4 Influence of CaO reactant on HCl release and control from inorganic and organic sources of chlorine 96 Release of Cl-content contaminants form NaCl is also investigated using a fixed bed tubular furnace. The experimental set-up as well as the experimental procedure and the sampling method are described in Section 2.2 & 2.3.

Figure 4 . 5 Figure 4 . 5

 4545 Figure 4. 5 Conversion of Cl -ion from NaCl under N 2 atmosphere

Figure 4 . 5

 45 Figure 4.5 from the tubular furnace results, increasing the temperature facilitates the evaporation of NaCl from the furnace. At temperature of 550 °C, no Cl -ion is detected from the absorption solution, indicating that the evaporation or conversion of NaCl is negligible.

Figure 4 . 6

 46 Figure 4. 6 Conversion of Cl -ion from NaCl under air atmosphere and O 2 atmosphere and the effect of water and SiO 2

  experiments of PVC are performed under N 2 atmosphere at heating rate of 10 o C/min. The experimental results are exhibited in Figure 4.7.

Figure 4 . 7

 47 Figure 4. 7 TG, DTG and 2 nd order DTG curves of PVC pyrolysis under N 2 atmosphere at heating rate of 10 o C/min

Figure 4 . 8

 48 Figure 4. 8 Products identification of the first mass loss peak during PVC pyrolysis using TG-FTIR analysis

Figure 4 . 8

 48 illustrates the identification results of the pyrolytic products of the maximum peak intensity in the first degradation stage. Results confirm that HCl generated from dehydrochlorination reaction of PVC is the most important product in the first degradation stage; meanwhile, a small amount of benzene is also formed[12]. According to Bockhorn et al. and Sørum et al., benzene is discharged immediately after HCl release from PVC since conjugated double bonds are

Chapter 4 : 101 4. 2 . 2 . 2 Figure 4 . 9 9 .

 4101222499 Figure 4. 9 HCl release from PVC pyrolysis and incineration in fixed bed tubular furnace reactor

2 100% 103 4. 3 . 1 Figure 4 . 4 . 10 .

 2103314410 Figure 4. 10 Effect of Ca/Cl molar ratio on HCl mitigation efficiency and CaO conversion rate during PVC pyrolysis with the presence of in-furnace CaO

Figure 4 . 1 .Figure 4 .

 414 Figure 4. 11 presents the SEM and EDS analysis of the spent CaO reactants at 750 o C with Ca/Cl molar ratio of 1:1 and 2:1. Besides, the elemental composition of the reactant surface obtained from EDS micro analyses are listed in Table 4.1.

4. 3 . 2 Figure 4 .Figure 4 . 12 .

 324412 Figure 4. 12 Effect of temperature on HCl mitigation efficiency during PVC pyrolysis with the presence of in-furnace CaO reactant

Figure 4 . 13 . 108 (

 413108 photograph, resulting in the formation of a very smooth and compact surface. However, at temperature of 850 o C, the surface becomes again very loose and porous, mainly attributed to the evaporation and decomposition of the product-layer, i.e. CaCl 2 . The SEM photographs confirm that at higher temperature, calcination and sintering of the reactant and the product-layer changes the surface properties and pore structure of the sorbent, thus reducing the capacity of CaO reactant for HCl mitigation.

Figure 4 .

 4 Figure 4. 13 SEM and EDS analysis of the spent CaO reactants at Ca/Cl molar ratio of 2:1: (1) 550 o C; (2) 650 o C; (3) 750 o C; (4) 850 o C; (e) SEM photograph of fresh CaO reactant

Figure 4 . 13 .

 413 Figure 4. 13. Results reveal that with the increase of temperature, the relative intensity of elemental chlorine is significantly decreased. The molar fraction of chlorine of the selected surface is decreased from 20.4% at 550 o C to 4.0% at 850 o C. Table 4.3 summarizes the elemental composition of the selected area on the reactant surface by EDS analysis, which further confirms the decline of HCl mitigation efficiency with respect to the reduction of the chlorine contained in the selected surface.

Figure 4 .

 4 Figure 4. 14 Effect of CaO particle size on HCl mitigation efficiency during PVC pyrolysis with the presence of in-furnace CaO reactant

Figure 4 . 7 in 4 . 2 ,

 4742 Figure 4.7 in Section 4.2, there are two distinct mass loss peaks of PVC pyrolysis TG curves, among which the first mass loss peak corresponds almost exclusively to dehydrochlorination

Chapter 4 :

 4 Influence of CaO reactant on HCl release and control from inorganic and organic sources of chlorine 113 and the total number of the pseudo-components. The effects of the variation of the parameters H p , T p , W hf , and A s are depicted in Figure 4.15.

Figure 4 .

 4 Figure 4. 15 Effect of the parameters in the Fraser-Suzuki function: 1, the height, H p ; 2, the Peak temperature, T p ; 3, the half width, w hf ; and 4, the symmetry, A s

Figure 4 . 4 .Figure 4 .

 444 Figure 4. 16 Deconvoluted DTG curves of PVC and PVC with CaO reactant pyrolysis at 5 o C/min, 10 o C /min, and 20 o C /min

Figure 4 .

 4 Figure 4. 17 Kinetic plots for the pseudo-components of PVC pyrolysis between conversionrates of 0.1 and 0.9 using both OFW and KAS methods

Figure 4 .

 4 Figure 4. 18 Kinetic plots for the pseudo-components of PVC pyrolysis with CaO reactantbetween conversion rates of 0.1 and 0.9 using both OFW and KAS methods

Chapter 5 :

 5 Influence of CaO reactant on pyrolysis tar upgrading of waste single-component temperature on pyrolytic tar cracking and composition are experimentally investigated using Py-GC/MS apparatus. Pyrolytic tar from typical MSW single components, such as food waste (occupies the largest proportions in many countries and regions), wood biomass (represents typical biomass waste), and PVC (represents plastic) are identified.

Figure 5 . 1

 51 Figure 5. 1 Schematic diagram of in-situ and ex-situ of tar cracking during MSW pyrolysis, data source from Kabir et al.[7]

Figure 5 . 2 and

 52 Py-GC/MS with/without CaO reactant addition. Due to the principle of Py-GC/MS experimental set-up, only volatile organic compounds are able to be identified. Based on the gas chromatogram, Figure 5.3 present the identification results of pyrolytic tar composition from wood biomass at 750 o C with/without the addition of CaO reactant. A total of 39 and 40 compounds are identified from pyrolytic tar without CaO and with CaO, respectively. However, in order to simplify the presentation of the results, the most typical compounds as suggested by the gas chromatogram (the most obvious peaks) are presented in the Figures. The selected compounds represent more than 70% of the total peak area.

Figure 5 . 2 Figure 5 . 3 Figure 5 . 2 and 5 . 3 ,

 52535253 Figure 5. 2 Identified tar composition from wood biomass pyrolysis at 750 o C in the absence of CaO reactant

Figure 5 . 4 2 )Figure 5 . 4 .

 54254 Figure 5. 4 Composition of pyrolytic tar derived from wood biomass pyrolysis at 750 o C without/with the addition of CaO reactant (others here represent the detected compounds which are not presented in Figure 5.1 and Figure 5.2)

Figure 5 . 5

 55 Figure 5. 5 The relative content of BTXN and Phenols in pyrolytic tar derived from wood biomass pyrolysis at 750 o C without/with CaO reactant

Figure 5 . 5

 55 Figure5.5 presents the effect of CaO reactant on the relative content of BTXN and phenols in pyrolytic tar from wood biomass pyrolysis at 750 o C. As depicted, the addition of CaO reactant promotes the total yields of BTXN by 14.9%. Among the BTXN, the relative content of toluene, xylene and naphthalene is increased by 24.3%, 211.8%, and 134.5%, respectively, when CaO reactant is used; in contrast, benzene yield is declined by 23.1% as a result of CaO addition. The results agree well with previous conclusions indicating that the use of CaO reactant promoted the formation of aromatic compounds but had a different effect on BTXN;

Figure 5 . 6

 56 Figure 5. 6 Representative lignin pyrolysis, decarboxylation and decarbonylation of pyrolytic tar over CaO reactant

Figure 5 . 7 Figure 5 . 8

 5758 Figure 5. 7 Composition of pyrolytic tar derived from food waste pyrolysis at 750 o C without/with the addition of CaO reactant (others represent mainly nitrogen-content compounds)

Figure 5 . 8 Figure 5 . 9 Figure 5 .

 58595 Figure 5. 8 The relative content of BTX, phenols, and ketones in pyrolytic tar derived from food waste pyrolysis at 750 o C without/with CaO reactant

Figure 5 .

 5 Figure 5. 11 Chlorine-content compounds formation from PVC pyrolysis, data derived from Ruizhi ZHANG[22]

Figure 5 . 12 .Chapter 4 ,

 5124 As revealed, the majority of the detected compounds belongs to C6 -C10 compounds, sharing approximately 61% of the total peak area form both non-catalytic and catalytic cases. The addition of CaO reactant slightly increases the heavy compounds and lowers the relative content of C1 -C5 by 30.2%. As aforementioned in PVC is the most important organic chlorine source and the degradation of PVC is always accompanied with the release of HCl. Actually in this experiment, HCl release is recorded from non-catalytic experiment with a relative content of 3.18%. However, with the addition of CaO reactant, HCl is eliminated from the vapors. The results further confirm the role of CaO reactant on HCl gas removal.

Figure 5 .

 5 Figure 5. 12 Distributions of detected species of pyrolytic tar derived from PVC pyrolysis at

Figure 5 .Chapter 5 : 5 . 3 Figure 5 .Figure 5 .Chapter 5 :Figure 5 .Figure 5 .Figure 5 .Figure 5 .Chapter 5 :Figure 5 . 14 -Figure 5 . 10 )

 555355555555514510 Figure 5. 13 The relative content of BTX, Naphthalene, Indene, Biphenyl, Fluorene, and Phenanthrene in pyrolytic tar derived from PVC pyrolysis at 750 o C without/with CaO reactant

Figure 5 .Figure 5 .

 55 Figure 5. 20 Composition of pyrolytic tar derived from PVC pyrolysis at 550, 650, 750, and 850 o C with/without CaO addition (Others here represent CO 2 , HCl, and acid compounds)

Figure 5 .Figure 5 .

 55 Figure 5. 21 Distributions of detected species of pyrolytic tar derived from PVC pyrolysis as a function of carbon atom numbers

Figure 5 .

 5 Figure 5. 22 The relative content of BTX in pyrolytic tar derived from PVC pyrolysis without/with CaO reactant

Figure 5 .Figure 5 .

 55 Figure 5. 23 The relative content of naphthalenes and naphthalene in pyrolytic tar derived from PVC pyrolysis without/with CaO reactant

Figure 5 .Figure 5 .

 55 Figure 5. 24 Schematic diagram of the formation and cracking of naphthalene with CaO reactant



  The addition of CaO reactant promotes the formation of aromatic compounds and inhibits the relative content of oxygenated compounds in pyrolytic tar derived from wood biomass and food waste; Pyrolytic tar derived from PVC pyrolysis are mainly aromatic compounds, mainly due to the generation of abundant conjugated double bonds as a result of dehydrochlorination from PVC; and the use of CaO reactant eliminates the produced HCl from PVC pyrolytic tar species; High value added tar species, i.e. BTXN (benzene, toluene, xylene and naphthalene) are promoted by CaO reactant from wood biomass and food waste pyrolysis; however, the use of CaO reactant declines the relative content of BTX in pyrolytic tar from PVC;  The relative content of BTX and benzene is decreased monotonically as the increase of temperature from 550 to 850 o C and with the addition of CaO reactant from PVC pyrolysis; whereas the relative content of naphthalene is promoted simultaneously.

II. Section 6 . 3

 63 investigates the effect of in-furnace CaO reactant on released HCl mitigation. HCl mitigation efficiency under varied temperature and reactant properties (varied Ca/Cl molar ratio and CaO particle size) are examined and morphology analysis of the spent in-furnace CaO reactant are conducted; III. Section 6.4 focuses on the effect of temperature and in-furnace CaO reactant on S-MSW pyrolytic tar cracking and tar composition upgrading.

Figure 6 . 1 Figure 6 . 1 (

 6161 Figure 6. 1 Effect of temperature and CaO reactant on produced gas characteristics: a, total gas yield; and b, combustible gas composition (the carrier gas N 2 is subtracted from the produced gas)

Figure 6 . 1 (

 61 Figure 6.1 (b) presents the effect of temperature and CaO reactant on the produced gas composition. Pyrolysis temperature directly determines the pyrolytic gas composition by affecting waste decomposition reaction and gas-reforming reactions.Table 6.2 summarizes

Figure 6 . 2

 62 Figure 6. 2 Effect of temperature and CaO reactant on hydrogen yield (in mL/g-MSW)

CaO on HCl mitigation 6 . 3 . 1

 631 XRD analysis of the spent CaO reactant

Figure 6 . 3 Figure 6 . 3

 6363 Figure 6. 3 XRD spectrums of the obtained solid residues (from bottom to top): 650 o C with CaO reactant; 750 o C with CaO reactant; and 850 o C with CaO reactant

Figure 6 . 4

 64 Figure 6. 4 Effect of temperature and Ca/Cl molar ratio on HCl mitigation efficiency during

Figure 6 . 4

 64 illustrates the effect of temperature Chapter 6: Influence of CaO reactant on pyrolysis of simulated MSW: HCl mitigation and tar upgrading 163 and Ca/Cl molar ratio on HCl mitigation efficiency. The obtained results are in good agreement with that obtained from Section 4.3 using only PVC as the feedstock. Increasing the temperature causes adverse effects on HCl mitigation efficiency using CaO reactant. As depicted, for the conditions of Ca/Cl molar ratio equal to 2:1, the removal efficiency declines continuously from 92.4 to 65.2% when the temperature is increased from 550 to 850 o C. It is noted that higher temperature accelerates the decrease of HCl mitigation efficiency; the efficiency decreased by 8.8% from temperature of 550 to 750 o C, but reduced sharply by 18.3% when further increasing the temperature to 850 o C. According to the thermodynamic calculation conducted in Section 3.2, the binding capacity of CaO and HCl is limited at higher temperature due to the shift of the thermodynamic equilibrium of CaO dechlorination reactions (i.e. Eq. (6.8)). Results from XRD spectrums and SEM photographs also reveals that the CaClOH content in the solids reduced apparently at 850 o C due to its decomposition and evaporation. Therefore, the feasible temperature window for HCl mitigation under the experimental conditions is below 750 o C.

Figure 6 . 5

 65 Figure 6.5 shows the SEM photographs of raw CaO and obtained solid residues. As presented, the fresh CaO reactant is a low-porosity material. The SEM image of the solid residues obtained at temperature of 650 o C shows clearly the formation of porous solid products on the surface of CaO reactant, indicating that the chemical reaction of CaO reactant and pyrolysis products occurs. When increasing the reaction temperature to 750 o C, the microstructures of the reactant and surface products become quite smooth and dense, mainly attributed to the melting of the formed products (the melting temperature of CaCl 2 is

Figure 6 . 5 6 . 4 6 . 4 . 1 Figure 6 . 6 Figure 6 . 6

 65646416666 Figure 6. 5 SEM images of fresh (a) and spent CaO reactants (b, pyrolysis residues at 650 o C; c, pyrolysis residues at 750 o C; and d, pyrolysis residues at 850 o C)

Figure 6 . 6 (

 66 a), tar content declines largely from 0.47 to 0.19 g/g-S-MSW as the pyrolysis temperature increased from 550 to 850 o C. Increasing the temperature significantly favors the decomposition of large organic compounds to produce light hydrocarbons as well as incondensable gases. When in-furnace CaO reactant is used, tar catalytic thermal cracking reactions are accelerated and tar yield is thus reduced and ranges from 0.38 to 0.14 g/g-S-MSW. With respect to tar reduction efficiencies as presented in Figure 6.6 (b), tar thermal reduction efficiency is significantly increased from 0 to 60.3% when increasing the temperature from 550 to 850 o C. When CaO is introduced into the furnace, tar catalytic thermal reduction efficiency is dramatically promoted from 19.2 to 70.1% under the experimental temperature. It is worth mentioning that the increment of tar catalytic thermal reduction efficiency declines with the increase of temperature. With the addition of CaO, tar catalytic thermal efficiency raises dramatically by 48.3% when pyrolytic temperature is increased from 550 to 750 o C; however, further increase of the temperature to 850 o C results

Chapter 6 : 6 . 4 . 2 Figure 6 . 7 Figure 6 . 8 Figure 6 . 9 Figure 6 . 7 ,Figure 6 . 8 and

 66426768696768 Figure 6. 7 Tar produced form S-MSW pyrolysis without CaO reactant at 550 o C

Figure 6 . 7 ,Figure 6 . 8 and 6 . 9 ,

 676869 Figure at lower temperature of 550 o C, aliphatic hydrocarbons are the major compounds in tar and only few aromatic species are detected. It is mainly due to the relative high content of plastics (PE 15% and PVC 10%) in S-MSW. Plastics easily decompose at low temperature and produce a great amount of olefins.

Figure 6 .Figure 6 .

 66 Figure 6. 10 Distributions of products in pyrolytic tar with and without the addition of CaO reactant: (a) pyrolysis at 550 o C and (b) pyrolysis at 750 o C

Figure 7 . 1 .

 71 S1represent conventional waste incineration plants and it is set as the baseline scenario for the comparison; S2 is a pyrolysis-based WtE system; S3 represent gasification with ex-situ high temperature dehydrochlorination as performed by Lahti Energia; and S4 refers to gasification-melting technology.

Figure 7 . 1 7 . 1 .

 7171 Figure 7. 1 System boundary of the considered scenarios

Figure 7 . 2

 72 Figure 7. 2 Schematic diagram of the considered incineration system

5 to 8 %

 8 , at temperature of approximately 1,200 °C. The residues are considered to be landfilled and the metals are recycled. The plant electricity efficiency (gross) is around 25.9%. The same APC systems as incineration systems are considered.

Figure 7 . 3 Figure 7 . 4

 7374 Figure 7. 3 Schematic diagram of the considered pyrolysis system, data obtained from the report of DGEngineering[29]

Chapter 7 :

 7 Life cycle assessment of MSW thermal technologies with different dehydrochlorination options 184 world. The MSW is charged into a shaft-furnace type gasifier from the top with coke and limestone, and the ash is melt at the bottom by O 2 -rich air at 1000-1800 °C. No

Figure 7 . 5

 75 Figure 7. 5 Schematic diagram of the considered gasification-melting system, data obtained from the work of Tanigaki et al.[33]

Chapter 7 :

 7 Life cycle assessment of MSW thermal technologies with different dehydrochlorination options185

Figure 7 . 6

 76 Figure 7. 6 reports the characterized environmental impacts of the considered four MSW thermal conversion systems. As aforementioned, the environmental impact categories of three non-toxic (i.e. GWP, AP, and EP), and three toxic (i.e. TETP, FAETP, and HTP) parameters are taken into consideration in this work. Results reveal that negative impacts (i.e.

7. 2 . 5 . 2 Figure 7 . 7

 25277 Figure 7. 7 Contribution analysis of the selected environmental impact categories

Figure 7 . 7 reports

 77 the contribution analysis results of different contributors. Generally, the impacts f o sta k e issio s , e e g e o e a d etals e li g a e the ost ote o th contributors, with the former one being the most important environmental burdens and the latter two being environmental savings (negative values as shown in Figure 7. 7). E e g e o e , due to the su stitutio of fossil fuels o ustio elated to ele t i it p odu tio , akes a sig ifi a t o t i utio to edu e the GWP, AP, a d EP i pa ts. Metals e li g

Figure 7 . 8 .Chapter 7 :Figure 7 . 8

 78778 Figure 7. 8 The considered systems focusing on different de-HCl options (a) and system boundaries (b)

Chapter 7 :

 7 Life cycle assessment of MSW thermal technologies with different dehydrochlorination options196 d due to the addition of CaO, the amount of bottom ash is increased. The calculation method includes the complete absorption of Cl by CaO reactant and CO 2 absorption by CaO;

Figure 7 . 9 Figure 7 . 9

 7979 Figure 7. 9 Contribution analysis of the selected environmental impact categories from the o side ed s ste s: S ', the p oposed h potheti al gasifi atio -based system; S3 gasification system; and S4, gasification-melting system



  mg/g-CaO with the increase of temperature from 550 to 850 o C. Higher raw HCl concentration increase the breakthrough and saturation time of CaO bed, but decrease HCl mitigation capacities due to the reduction of both physical adsorption and chemical reaction rates. The presence of CO and H 2 will occupy the number of free active sites on CaO surface and cause adverse effects on HCl mitigation.A kinetics analysis has been developed by comparing experimental data with models describing the reaction at the particle surface. It has been concluded that the HCl mitigation at the CaO particle surface is firstly limited by the heterogeneous gas-solid reaction, and then by the HCl diffusion through this porous growing layer. Furthermore, morphology analysis of the spent CaO reactants are examined to reveal the formation process of the product-layer; and the spent CaO physical properties further indicates that both the shifts of chemical equilibrium and the reduction of surface area and closure of pore structures at higher temperature are responsible for the decline of HCl mitigation efficiency by CaO. HCl release from organic and inorganic sources and effect of in-furnace CaO reactant on HCl mitigation and pyrolytic tar upgradingTo simulate the in-situ generation of HCl from organic source, pyro-gasification of PVC has been performed with or without CaO addition. The experimental data have been then used for the modeling of the different PVC decomposition steps. Although the average apparent activation energy of pseudo dehydrochlorination reaction is increased from 136.52 to 152.61 kJ/mol with the addition of CaO, the apparent activation energy of the overall PVC decomposition has been decreased from 197.26 to 148.92 kJ/mol using CaO reactant.HCl emission from organic and inorganic sources has also been conducted using simulated MSW. Results reveal that in-situ addition of CaO reactant is effective for HCl mitigation. CaClOH is identified as the main dehydrochlorination product by XRD observations. Increasing the temperature causes adverse effects on HCl mitigation efficiency and the increase of Ca/Cl molar ratio promotes the efficiency. The optimized working conditions observed in this work are at temperature of 750 o C and Ca/Cl molar ratio of 2:1.The role of CaO reactant on pyrolytic tar is also determined by identifying the chemical speciation of the produced tars under catalytic and non-catalytic conditions. Using CaO, the amount of oxygenated organic compounds has been reduced, improving the quality of the tars for a further bio-oil upgrading.Life cycle assessment of pyro-gasification-based WtE systems and WtE systems with different dehydrochlorination optionsFirstly, LCA investigation of four WtE thermal conversion systems (including S1, conventional waste incineration system; S2, pyrolysis-based WtE system; S3, gasification system with ex-situ dehydrochlorination as performed by Lahti Energia; and S4, gasification-melting technology) is conducted to determine the environmental performance of pyro-gasification-based WtE systems when compared with conventional MSW incineration.

Figure 3 . 0 . 1

 301 Figure 3.0.1 Mécanisme de réaction entre CaO et HCl proposé : le premier régime correspond à une limitation de la cinétique par la réaction hétérogène de surface, laquelle conduit à la formation de CaCl 2 (en gris clair). Plus la réaction avance, plus le réactif CaO est converti en CaCl 2 ui ep se te u e ou he s' paississa t à la su fa e du CaO ui 'a pas encore réagi. Ceci traduit que la diffusio de HCl à l'i t ieu de ette ou he ai si formée devie t l' tape li ita te du p o essus. L'aug e tatio de de sit de la ou he, observée notamment par microscopie électronique à balayage, bloque la réaction en inhibant la diffusion.

  systèmes de traitement thermique des déchets municipaux existants : incinération conventionnelle (S1), pyrolyse (S2), gazéification (S3), et le système de gazéification / fusion des e d es " . L ́tude se o e t e su l i pa t e i o e e tal de ha ue s stè e et les p i ipau p o essus e ause. Le ṕ i èt e d ́tude est e poś à la figu e . . .

Figure 7 . 0 .

 70 Figure 7.0. P i t e d' tude des 4 scénarios évalués: S1 = incinération, S2 = pyrolyse, S3 = gazéification et S4 = gazéification et fusion de cendres

Figure 7 . 0 . 2 Résumé du Chapitre 8 Conclusion

 7028 Figure 7.0.2 Présentation des 3 scénarios couplant le traitement thermique des sidus solides ave l'a atte e t du hlo e, pa thode e -situ ou in-situ : S4, gazéification avec fusion des cendres puis traitement du chlore ex-situ à basse température (APC) ; S3, gazéification puis épuration du chlore ex-situ à haute température ; et S ', gazéification et abattement du chlore in-situ en présence du réactif CaO

Figure 1 . 1 2 Figure 1 . 2 3 Figure 1 . 3 4 Figure 1 . 4 5 Figure 1 . 5 6 Figure 1 . 6 9 Figure 1 . 7 10 Figure 1 . 8 16 Figure 1 . 9 21 Figure 1 . 22 Figure 1 .

 1121231341451561691710181619211221 Figure 1. 1 MSW generation in France and China during 2007-2010 ........................... 2 Figure 1. 2 Preference and contents of the waste hierarchy as regulated by European Directive, data derived from [6] ................................................................................ 3 Figure 1. 3 Share of MSW treatment methods in the EU during 2004-2015, data source from Eurostat database [7] ....................................................................................... 4 Figure 1. 4 Conceptual diagram of Kymijärvi II RDF gasification power plant, located in Lahti region, Finland, data acquired from [13] ......................................................... 5 Figure 1. 5 The main ideas of this study: investigate in-furnace capture of chlorine to avoid the release of HCl by using in-situ CaO reactant ............................................ 6 Figure 1. 6 Graphic principle of acid rain formation pathway: (1) Emissions of SO 2 , NO x and HCl are released into the air, where (2) the pollutants are transformed into acid particles that may be transported in a long distance, (3) the acid particles then fell to the earth as wet and dry deposition (dust, rain, snow, or etc.) and (4) may cause harmful effects on soil, forests and water, data obtained from [42] .... 9 Figure 1. 7 Schematic diagram of high temperature corrosion occurred on the surface of the equipment, data obtained from Wang [49] ................................................. 10 Figure 1. 8 Spray Drying Gas Cleaning System, data obtained from Razvan Lisnic et al. [74] ........................................................................................................................... 16 Figure 1. 9 Phases and applications of an LCA standardized by ISO, data obtained from the work of Jun Dong [102] ..................................................................................... 21 Figure 1. 10 LCA concerns in waste management field, data source from Laurent et al. [125] ......................................................................................................................... 22 Figure 1. 11 Structure and main contents of the thesis ................................................. 24

Figure 2 . 1 38 Figure 2 . 2 40 Figure 2 . 3 N 41 Figure 2 . 4 43 Figure 2 . 5 46 Figure 2 . 6 48 Figure 2 . 7

 21382240234124432546264827 Figure 2. 1 Pictures of MSW components: (a), PVC pellet; (b), wood; (c), food waste; (d), PE pellet; and (e), paper ................................................................................... 38 Figure 2. 2 XRD analysis of CaO reactant ........................................................................ 40 Figure 2. 3 N 2 adsorption-desorption (a) isotherms and (b) Log isotherms of CaO reactant with particle size of 0.25 -0.43 mm (40-60 mesh) .................................. 41 Figure 2. 4 General photograph and schematic diagram of the vertical tubular furnace reactor: 1, gas source; 2, gas flow meter; 3, vertical quartz tubular reactor; 4, electric heater and temperature controller; 5, quartz filter; 6, CaO reactant for HCl mitigation; 7, quartz wool; 8, HCl gas detector ...................................................... 43 Figure 2. 5 Schematic diagram of the experimental fixed bed tubular furnace apparatus. 1, gas source; 2, gas flowmeter; 3, quartz tubular furnace; 4, quartz boat; 5, electric heater and temperature controller; 6, condenser for tar collection; 7, cooling water; 8, HCl mitigation solution (0.1 mol/L NaOH); 9, silica gel; 10, ice-water bath ......................................................................................................... 46Figure 2. 6 General photograph and schematic diagram of the Py-GC/MS experimental apparatus ................................................................................................................. 48Figure 2. 7 General principle of the thermodynamic equilibrium calculation system .. 54

Figure 2 . 8 56 Figure 3 . 1 66 Figure 3 . 2 67 Figure 3 . 3 68 Figure 3 . 4 69 Figure 3 . 5 Normalized reaction time as a function of CaO reactant conversion rate 70 Figure 3 . 6 73 Figure 3 . 7 Figure 3 . 8 75 Figure 3 . 9 75 Figure 3 . 10 77 Figure 3 . 11 77 Figure 3 . 12 79 Figure 3 . 13 79 Figure 3 . 81 Figure 3 . 15 83 Figure 4 . 1

 28563166326733683469357036733738753975310773117731279313793813158341 Figure 2. 8 Main menu of FactSage 6.3 software ........................................................... 56

Figure 4 . 2 92 Figure 4 . 3 93 Figure 4 . 4 95 Figure 4 . 5 96 Figure 4 . 6

 429243934495459646 Figure 4. 2 Thermodynamic simulation results of pure NaCl under N 2 atmosphere .... 92 Figure 4. 3 Thermodynamic simulation results of NaCl under considered working conditions: (a) under O 2 atmosphere; (b) under N 2 atmosphere with water; (c) under O 2 atmosphere with water; and (d) with the presence of SiO 2 and water under O 2 atmosphere .............................................................................................. 93 Figure 4. 4 TG analysis results of NaCl under N 2 atmosphere ........................................ 95 Figure 4. 5 Conversion of Cl -ion from NaCl under N 2 atmosphere ................................ 96Figure 4. 6 Conversion of Cl -ion from NaCl under air atmosphere and O 2 atmosphere and the effect of water and SiO 2 ............................................................................. 97

Figure 4 . 7

 47 Figure 4. 7 TG, DTG and 2 nd order DTG curves of PVC pyrolysis under N 2 atmosphere at heating rate of 10 o C/min ........................................................................................ 98

Figure 4 . 8 100 Figure 4 . 9

 4810049 Figure 4. 8 Products identification of the first mass loss peak during PVC pyrolysis using TG-FTIR analysis ........................................................................................... 100 Figure 4. 9 HCl release from PVC pyrolysis and incineration in fixed bed tubular furnace reactor ................................................................................................................... 101

Figure 4 . 103 Figure 4 . 4 ) 105 Figure 4 . 12 106 Figure 4 . 2 : 1 : ( 1 ) 108 Figure 4 . 14 109 Figure 4 . 113 Figure 4 .

 410344105412106421110841410941134 Figure 4. 10 Effect of Ca/Cl molar ratio on HCl mitigation efficiency and CaO conversion rate during PVC pyrolysis with the presence of in-furnace CaO .......................... 103 Figure 4. 11 SEM and EDS detection of the spent CaO reactants: (1) SEM, Ca/Cl molar ratio: 1:1, 750 o C; (2) SEM, Ca/Cl molar ratio: 2:1, 750 o C; (3) EDS, Ca/Cl molar ratio: 1:1, 750 o C; (4) EDS, Ca/Cl molar ratio: 2:1, 750 o C ..................................... 105 Figure 4. 12 Effect of temperature on HCl mitigation efficiency during PVC pyrolysis with the presence of in-furnace CaO reactant ..................................................... 106 Figure 4. 13 SEM and EDS analysis of the spent CaO reactants at Ca/Cl molar ratio of 2:1: (1) 550 o C; (2) 650 o C; (3) 750 o C; (4) 850 o C; (e) SEM photograph of fresh CaO reactant .................................................................................................................. 108 Figure 4. 14 Effect of CaO particle size on HCl mitigation efficiency during PVC pyrolysis with the presence of in-furnace CaO reactant ..................................................... 109Figure 4. 15 Effect of the parameters in the Fraser-Suzuki function: 1, the height, H p ; 2, the Peak temperature, T p ; 3, the half width, w hf ; and 4, the symmetry, A s ........ 113Figure 4. 16 Deconvoluted DTG curves of PVC and PVC with CaO reactant pyrolysis at 5 o C/min, 10 o C /min, and 20 o C /min ....................................................................... 115

Figure 4 .

 4 Figure 4. 17 Kinetic plots for the pseudo-components of PVC pyrolysis between conversion rates of 0.1 and 0.9 using both OFW and KAS methods ................... 117

Figure 4 . 118 Figure 5 . 1 Figure 5 . 2 Figure 5 . 3 Figure 5 . 4 Figure 5 . 5 Figure 5 . 6 Figure 5 . 7 Figure 5 . 8 Figure 5 . 9 Figure 5 . 10 Figure 5 Figure 5 .Figure 5 . 13 Figure 5 . 14 Figure 5 . 15 Figure 5 .Figure 5 .Figure 5 . 18 Figure 5 . 19 Figure 5 . 20 Figure 5 .Figure 5 . 22 Figure 5 . 23 Figure 5 . 24 Figure 6 . 1 Figure 6 . 2 Figure 6 . 3 Figure 6 . 4 Figure 6 . 5 Figure 6 . 6 Figure 6 . 7 Figure 6 . 8 Figure 6 . 9 Figure 6 .Figure 7 . 1 Figure 7 . 2 Figure 7 . 3 Figure 7 . 4 Figure 7 . 5

 41185152535455565758595105551351451555518519520552252352461626364656667686967172737475 Figure 4. 18 Kinetic plots for the pseudo-components of PVC pyrolysis with CaO reactant between conversion rates of 0.1 and 0.9 using both OFW and KAS methods ................................................................................................................. 118

Table 1 . 1 Chemical reactions related to high temperature corrosion

 1

							Chemical reaction	Equation
						Fe		2	HCl		FeCl	2		H	2	1.8
						FeO		2	HCl		2 FeCl H O 2 	1.9
	2 Fe O 3		2	HCl CO 		FeO FeCl 	2		2 H O CO 2 	1.10
	3 Fe O 4		2	HCl CO 		2	FeO FeCl 	2		2 H O CO 2 	1.11
		2 Cr O 3		4	HCl H 	2		2 CrCl	2		3	2 H O	1.12
							2	Fe			3 Cl	2		2	FeCl	3	1.13
										Cr Cl 	2		CrCl	2	1.14
				2 Cr O 3 2		4 Cl	2		2 O		2 CrO Cl 4	2

Table 1 . 2 HCl emission standards for MSW incineration plants in China, the European Union and the United States

 1

	[55-58]

Table 1 . 3 advantages and disadvantages of the aforementioned three HCl removal methods

 1

	HCl
	removal	Advantages	Disadvantages
	method
			1, excessive use of sorbent due to the low
			reaction rates (typically three times higher the
			stoichiometric amounts); 2, removal efficiency
			(typically at 50 to 80 %) is not high enough to
		Operation temperature is	meet the emission standard; 3, the efficiency of
		relatively high (thus spray	downstream particle removal equipment is
	Dry	de-superheating system is	affected due to the excessive use of fine
	treatment	not essential);
	method	Equipment is simple and
		inexpensive;
		No wastewater or corrosion.
		Relatively high removal
		efficiency, more than 95 %
	Wet treatment method	for HCl removal; Organic pollutants and heavy metals can be removed at the same time by cooling
		and using of scrubbing
		solution.

sorbents; 4, the requirement for the design and distribution of nozzles is improved, since the sorbents needs to be uniformly distributed throughout the entire reaction zone; and 5, costs are high due to the preparation of the sorbent (acquisition, purification, transportation, grind, storage and etc.).

Table 2 .1 andTable 2 .2.

 22

Table 2 . 1 Proximate and ultimate analysis of MSW components

 2

		M	Proximate analysis a (ad, wt. %) V FC	A	C	Ultimate analysis a (ad, wt. %) H O b N	S	Cl	LHV a (MJ/kg)
	PVC	0.00 95.80 4.20 0.00	38.90 4.90 0.00 0.00 0.00 56.20 19.30
	Wood 9.65 55.25 33.89 1.20	42.90 5.43 39.96 0.27 0.58 0.00	17.04
	Paper 7.64 71.61 7.66 13.09	34.05 3.24 41.82 0.12 0.04 0.00	12.81
	Bread 6.43 71.33 20.77 1.46	43.40 5.53 39.04 1.99 0.36 1.79	17.11
	PE	0.14 98.47 0.09 1.30	83.08 11.20 4.07 0.01 0.20 0.00	45.48

a M: moisture; V: volatiles; FC: fixed carbon; A: ash; C: carbon; H: hydrogen; O: oxygen; N: nitrogen; S: sulfur; LHV: lower heating value; b by difference.

Table 2 . 2 Overall composition and proximate and ultimate analysis of S-MSW

 2

				Overall composition (ar, wt.%) a
	Food waste		Paper	Polyethylene (PE)		Polyvinyl chloride (PVC)		Wood
	55.0		15.0	15.0			10.0			5.0
	Proximate analysis (ad, wt.%) a		Ultimate analysis (ad, wt.%) a
	Moisture	Ash	Volatile	Fixed Carbon	C	H	N	S	Cl	O b
	5.19	3.02	77.09	14.70	47.47 5.97 1.13 0.26 6.60 30.35

a according to the as-received basis; b by difference.

Table 2 . 3 Physicochemical properties of the prepared CaO reactant

 2

	Particle size,	Specific surface area,	Total pore volume,	Average pore diameter,
	mm	m 2 /g	cm 3 /g	nm
	0.25 -0.43 mm (40-60 mesh)	2.37	0.039	25.0
	0.13-0.18 mm (80-120 mesh)	6.28	0.097	23.1

Table 2 . 4 Use of feedstock and reactant in different experiments

 2

	Chapter	Experiments	Feedstock	Reactant	Apparatus
	3	HCl gas mitigation by CaO reactant	Gas mixtures	CaO	Vertical tubular furnace reactor
	4	HCl release and mitigation characteristics from inorganic source (NaCl) and organic source (PVC)	NaCl; PVC	CaO	TGA Fixed bed tubular furnace
		Tar formation and	MSW single
	5	composition during MSW pyrolysis with/without CaO	component: Food, PVC, and	CaO	Py-GC/MS
		reactant	Wood
	6	Characteristics of S-MSW pyrolysis with/without CaO	Simulated-MSW	CaO	Fixed bed tubular
		reactant

furnace 2.

2 Experimental apparatus and procedures 2.2.1 Reaction between CaO and HCl gas at moderate high temperature 2.2.1.1 Vertical tubular furnace reactor

Table 2.5

	Chapter 2: Materials and methods
	presents the experimental conditions. Temperature series at moderate high level of 550, 650,
	750, and 850 °C is selected; and the effect of raw HCl concentration (2000 ppmv, 1000 ppmv,
	and 500ppmv) is also considered. Besides, the effect of CO 2 content (12 vol. %) as well as the
	influence of simulated syngas (the simulated syngas contains 1.6 vol. % H 2 , 1.2 vol. % CH 4 , 2.5
	vol. % CO, 6 vol. % CO 2 , and N 2 balance; for these experiments, 50 vol. % 2000 ppmv HCl gas
	and 50 vol. % simulated syngas are used). Repeated runs are conducted under identical
	conditions to ensure the repeatability of the results.

Table 2 . 5 Experimental conditions of HCl mitigation by CaO reactant at moderate high temperature

 2

	No.	Temperature, o C	HCl concentration, ppmV	Gas flowrate, L/min	Atmosphere
	1#	550
	2# 3#	650 750	2000	2.0	N 2
	4#	850
	5# 6#	750	1000 500	2.0 2.0	N 2
	7#		1000	2.0	12 vol.% CO 2
	8#	750	1000	2.0	50 vol.% simulated syngas
	Note: the effect of temperature: 1#, 2#, 3# and 4#; the effect of raw HCl concentration: 3#,
	5# and 6#; the effect of atmosphere: 3#, 7# and 8#.
		Before the experiments, air leakage of the reactor and connection pipes have to be
	carefully tested, since the leakage of HCl and CO gases are quite dangerous. At the beginning
	of the experiments, 2.0 g of CaO reactant is accurately weighed and placed on the top of the
	quartz filter. Quartz wool is used to protect CaO reactant bed. Then the reactor is heated to a
	desired temperature with 2.0 L/min N 2 as purge gas. At the same time, the detector is turned
	on and connected for preheating and purging. When the temperature of the reactor as well
	as the HCl gas detector reaches the desired temperature and becomes stable, the inlet gas is
	switched to the experimental conditions. During the experiments, HCl concentration is
	monitored and recorded every 2 seconds by the detector. The outlet gas from the detector is
	cleaned by 0.1 mol/L NaOH solutions before being discharged.

Table 2

 2

	.6 illustrates the experimental conditions conducted in
	Chapter 4. The release characteristics of HCl from NaCl and PVC under redox atmosphere are
	identified. Besides, the effect of CaO reactant on HCl removal efficiency during PVC pyrolysis
	is studied. Temperature series of 550, 650, 750, and 850 o C, Ca/Cl molar ratio of 1:2, 1:1, 2:1,
	and 4:1, and CaO particle sizes of 0.25-0.43 mm and 0.13-0.18 mm are compared.

Table 2 . 6 Experimental conditions of the influence of CaO reactant on HCl release and control from inorganic and organic sources

 2

		Temperature,		Ca/Cl
	Feedstock		Reactant		Atmosphere
		o C		molar ratio
	NaCl	550, 650, 750, 850, 950	no	-	N 2 ; O 2 ; Air; Air + water vapor
		100, 150, 200, 225, 250,
	PVC	275, 300, 325, 350, 400,	no	-	N 2 ; Air
		500, 600, 700, 800
	PVC	550, 650, 750, 850	40-60 mesh CaO 80-120 mesh CaO	1:2; 1:1; 2:1; 4:1	N 2

set of experiments are also conducted in Chapter 4 by thermogravimetric analyzer

	(TGA, 951 Dupont Instrument) to record the mass change of non-catalytic and catalytic
	pyrolysis of PVC. The analyzer from Dupont Instrument is a horizontal module composed of
	two separate chambers disposed on the two sides of the balance housing. The experimental
	procedures of the TGA are the following. For non-catalytic PVC pyrolysis, pure PVC is
	accurately weighted and placed into the platinum pan of the analyzer; the apparatus is then
	heated from room temperature to 900 o C at different heating rate of 5, 10, and 20 o C/min
	under N 2 atmosphere. For catalytic pyrolysis cases, pure PVC is firstly mixed with CaO
	reactant at specific Ca/Cl molar ratio. Then the mixture is accurately weighted and placed
	into the analyzer; the same heating rate and temperature program are adopted as
	non-catalytic cases. The obtained results are then used for kinetics modeling of PVC
	dehydrochlorination process.

2.2.3 Influence of CaO reactant on Tar upgrading from Waste Single Component Pyrolysis 2.2.3.1 Pyrolysis-Gas Chromatography/Mass Spectrometry (Py-GC/MS) experimental apparatus

2.2.3.2 Operating conditions and experimental procedure Table 2.7 summarizes the experimental conditions conducted in Chapter 5. The effect of

	Chapter 2: Materials and methods
	investigated by Py-GC/MS apparatus. Besides, the effect of temperature and CaO reactant on
	PVC pyrolysis tar formation is further compared.
	0.25µm is used as the chromatographic
	separation; the purge gas is Helium (the purity is 99.999%) with flow rate of 1.0
	L/min; the temperature is programmed to increase from 40 o C to 250 o C at 8 o C/min;
	M/Z values between 35 to 450 are detected by the mass spectrometer at a scan rate
	of 2337.1 amu/s.
	CaO reactant on tar formation and tar composition from MSW single components are

Table 2 . 7 Experimental conditions of the effect of temperature and CaO reactant on tar cracking and upgrading

 2

	No.	Feedstock	Temperature, o C	Reactant	Reactor type
	1# 2# 3#	Food Wood PVC	750	Without CaO	Py-GC/MS apparatus
	4# 5# 6#	Food Wood PVC	750	With CaO	Py-GC/MS apparatus
	7# 8# 9#	PVC	550 650 850	Without CaO	Py-GC/MS apparatus
	10# 11# 12#	PVC	550 650 850	With CaO	Py-GC/MS apparatus
	Before the experiments, the pyrolyzer is successively filled with some quartz wool,
	feedstock, and further some quartz wool. For CaO reactant pyrolysis experiments, CaO is
	uniformly mixed with the feedstock prior to be added into the pyrolyzer. For each run,
	approximately 5 mg of feedstock is used. Fast pyrolysis analysis is performed under Helium
	atmosphere and the heating rate is set as 10 000 °C/s to the desired final temperature.

2.2.4 Influence of CaO reactant on simulated MSW pyrolysis

	Chapter 2: Materials and methods
	Simulated MSW, as aforementioned in Section 2.2.1, is used as feedstock to conduct
	pyrolysis experiments with/without CaO reactant in Chapter 6. The experimental apparatus
	used is fixed type tubular furnace as illustrated in Section 2.2.2. The experimental
	procedures are also similar; besides gas products from S-MSW pyrolysis are collected,
	identified and analyzed.

Table 2 . 8 Experimental conditions of influence of CaO reactant on simulated MSW pyrolysis

 2

	Experimental runs	Temperature oC	CaO reactant	molar ration Ca/Cl	a	b	c
	1#	550
	2# 3#	650 750	Without	-
	4#	850
	5#	550
	6# 7#	650 750	With	2:1
	8#	850
	9#			1:2
	10#	650	With	1:1
	11#			4:1
	12#			1:2
	13#	750	With	1:1
	14#			4:1
	a focusing on the effect of temperature on product yields and properties during S-MSW
	pyrolysis;
	b focusing on the effect of temperature and CaO reactant on product yields and
	properties during S-MSW pyrolysis;
	c focusing on the effect of temperature and Ca/Cl molar ratio on HCl mitigation.
	Before the test, tubular furnace is firstly preheated to the desired temperature by
	electrical heater under N 2 atmosphere. When the temperature is stable, 5.0 g of S-MSW is
	accurately weighed and placed in a dried quartz boat. For experiments with CaO reactant,
	S-MSW is uniformly mixed with CaO reactant before being added into the boat. Then N 2
	carrier gas flow is adjusted to 120 NmL/min and the quartz boat is pushed into the center of
	the quartz tubular furnace in a short time to avoid air leaking in. The reaction time is set

Table 2 . 9 Impact categories included in the method CML (baseline)

 2

	Method: CML (baseline)
	Impact category group	Name of the impact category in the method	Characterization unit
	Acidification Potential	Acidification potential -average Europe	kg SO 2 equivalent
	Climate change Potential Climate change -GWP100	kg CO 2 equivalent
	Depletion of abiotic	Depletion of abiotic resources -elements, ultimate reserves	kg antimony equivalent
	resources Potential	Depletion of abiotic resources -fossil fuels	MJ
		Freshwater aquatic ecotoxicity -	kg 1,4-dichlorobenzene
		FAETP inf	equivalent
	Ecotoxicity Potential	Marine aquatic ecotoxicity -MAETP inf	kg 1,4-dichlorobenzene equivalent
		Terrestrial ecotoxicity -TETP inf	kg 1,4-dichlorobenzene equivalent
	Eutrophication Potential	Eutrophication -generic	kg PO 4 -equivalent
	Human toxicity Potential	Human toxicity -HTP inf	kg 1,4-dichlorobenzene equivalent
	Ozone layer depletion Potential	Ozone layer depletion -ODP steady state	kg CFC-11 equivalent
	Photochemical oxidation Potential	Photochemical oxidation -high NOx	kg ethylene equivalent
	Among the categories, six of them are considered and compared in this work, including
	three non-toxic and three toxic impacts: non-toxic, Climate change -GWP100, Acidification
	Potential -AP, and Eutrophication Potential -EP; and toxic, Terrestrial ecotoxicity -TETP,
	Freshwater aquatic ecotoxicity -FAETP and Human toxicity -HTP. Detailed characterization
	unit and normalization references have been indicated

in Table 2.10.

	Chapter 2: Materials and methods

Table 2 . 10 Selected life cycle impact categories, characterization and normalization references based on CML (baseline) method from GaBi software

 2

					Normalization(in
	Impact categories	Unit	Characterization Factor	per year, world,
					2000)
	Global warming		CO 2	1
	potential	kg CO 2 -equivalent	CH 4	28	4.22E+13
	(GWP100)		N 2 O	265
			SO 2	1.2
	Acidification potential (AP)	kg SO 2 -equivalent	NO x HCl	0.5 0.88	2.39E+11
			NH 3	1.60
	Eutrophication Potential (EP)	kg PO 4--equivalent	NO x N 2 O NH 3	0.13 0.27 0.35	1.58E+11
	Terrestrial	kg	HF	2.9E-03
	ecotoxicity (TETP)	1,4-dichlorobenzene eq.	PCDD/Fs	1.2E+04	1.09E+12
	Freshwater	kg	HF	4.6
	aquatic ecotoxicity (FAETP)	1,4-dichlorobenzene eq.	PCDD/Fs	2.1E+06	2.36E+12
			HCl	0.5
	Human toxicity (HTP)	kg 1,4-dichlorobenzene eq.	PM NH 3 HF NOx	0.82 0.1 2.9E+03 1.2	2.58E+12
			SO2	9.6E-02

Chapter 3 Experimental and kinetics study on HCl mitigation by CaO reactant 3.1 Introduction

	As it has been indicated in Chapter 1, HCl is usually generated as a consequence of high
	chlorine (Cl) content in MSW. The presence of HCl in flue gas causes serious problems
	associated with corrosion, toxic organic contaminants formation, acidification, etc. Currently,
	HCl is usually removed together with acid gases by air pollution control (APC) systems in WtE
	plants, which basically works in the temperature range lower than 300 o C [1-3]. Although the
	removal efficiency of HCl by APC systems is high enough for the emission control of HCl gas,

Table 3 . 1 The specific heat capacity calculation coefficients and thermodynamic data

 3

	CaO	CaCl 2	HCl	H 2 O

Table 3 . 2 Thermodynamic calculation results of Gibbs free energy and standard

 3

			equilibrium constant
	T/K	Δ H K , kJ/ ol	ln K θ	K θ	Δ G K , kJ/ ol
	500	-216.41	38.78	6.97X10 16	-161.22
	550	-216.24	34.08	6.33 X10 14	-155.84
	600	-216.06	30.18	1.28 X10 13	-150.54
	650	-215.85	26.89	4.77 X10 11	-145.32
	700	-215.62	24.09	2.90 X10 10	-140.21
	750	-215.36	21.69	2.62 X10 9	-135.22
	800	-215.06	19.60	3.26 X10 8	-130.37
	850	-214.73	17.79	5.30 X10 7	-125.69
	900	-214.36	16.20	1.08 X10 7	-121.21
	950	-213.94	14.81	2.69 X10 6	-116.94
	1000	-213.48	13.58	7.93 X10 5	-112.93
	1050	-212.96	12.51	2.71 X10 5	-109.20
	1100	-212.39	11.57	1.06 X10 5	-105.80
	1150	-211.76	10.75	4.65 X10 4	-102.75
	1200	-211.07	10.03	2.28 X104	-100.11

Table 3 . 3 Estimation of kinetic parameters τ p , τ c , θ 2 , k s , D p , E c and E p of HCl mitigation by CaO at different temperature

 3

	Temperature, o C	τ c *10 5 s	τ p *10 5 s	θ 2	k s *10 -3 , m/s	D p *10 -8 , m 2 /s	E c , kJ/mol	E p , kJ/mol
	550	4.84	6.85	1.42	10.20	3.58
	650 750	2.53 1.46	5.34 5.16	2.11 3.5	5.91 3.42	4.60 4.34	44.17	17.39
	850	0.85	3.08	3.62	1.79	7.98

τ p and τ c represents the total time taken when the conversion is only controlled by product-layer diffusion o he i al ea tio , espe ti el ; θ 2 refers to the ratio of τ p /τ c ; k s and D p is reaction rate and effective diffusivity through the product-layer; E c and E p represent the apparent activation energy of chemical reaction and product-layer diffusion, respectively.

Table 3 . 4 Specific surface area, pore volume and porosity of the spent CaO reactant

 3

		Raw CaO	550 o C	650 o C	750 o C	850 o C
	S BET , m 2 /g	2.4	4.2	2.6	1.5	2.1
	Total pore volume*10 -2 , cm 3 /g	3.9	4.4	2.7	1.6	2.1
	Porosity, %	4.5	5.0	3.1	1.8	2.4

Table 3 . 5 Experimental conditions adopted to investigate the effect of gas composition on HCl mitigation by CaO reactant

 3

	No.	Temperature, o C	HCl concentration, ppm	Gas flowrate, L/min	Gas composition, vol.%
	1#	750	1000	2.0	CO 2 12, and N 2 balance
	2#	750	1000	2.0	H 2 1.6, CH 4 1.2, CO 2.5, CO 2 6.0, and N 2 balance

composition on HCl mitigation is further investigated. Experiments are conducted at temperature of 750 o C, gas flow rate of 2.0 L/min. The presence of 12% CO 2 and 50% simulated syngas are evaluated in the present study. Due to the introduction of CO 2 and simulated syngas, raw HCl concentration decreases to 1000 ppm, and the detailed gas compositions are summarized in Table

3

.5.

Table 3 . 6 Summary of HCl breakthrough experimental results

 3

							Maximum
	Fixed parameter	Varied parameter	Breakthrough time, min	Saturation time, min	HCl mitigation amount, mg/g-CaO	CaO bed conversion rate, %	HCl mitigation rate, mg/(g-CaO
							*min)
		550 o C	197.8	268.9	778.9	59.8	3.26
	2000 ppm;	650 o C	106.0	179.3	507.2	38.9	3.26
	2.0 L/min	750 o C	51.4	115.8	321.5	24.7	3.26
		850 o C	21.8	69.6	173.9	13.4	3.26
	2.0 L/min;	1000 ppm	105.9	200.0	264.9	20.3	1.63
	750 o C	500 ppm	177.1	326.1	215.4	16.5	0.81
	1000 ppm;	12% CO 2	87.9	209.98	274.5	21.05	1.63
	2.0 L/min;
	750 o C	50% syngas	74.2	164.3	218.5	16.8	1.63

  Chapter 4: Influence of CaO reactant on HCl release and control from inorganic and organic sources of chlorine[START_REF] Mura | On the kinetics of dry reaction between calcium oxide and gas hydrochloric acid[END_REF] To summarize, theoretically equilibrium simulation by using FactSage demonstrates that the thermodynamic behavior of NaCl under N 2 atmosphere is quite stable; no evaporation or thermochemical conversion of NaCl in the temperature range of 450 -950 o C is predicted.However, the interactions caused by O 2 atmosphere, water content, and the presence of SiO 2

	the total Na-compounds reaches 7.91%. Therefore, the corresponding chemical reaction, as
	previously mentioned in Section 1.2.1.1, is responsible for the high conversion of chlorine
	from NaCl to HCl, and can be described as:
	2	NaCl		2	2 SiO		2 H O		2	2 HCl Na Si O 2 5 	(4.2)

2 (g) remains at a same level at around 2.69% and 2.01%, respectively, if comparing those with the case of SiO 2 addition. However, the molar fraction of HCl increases largely as compared to the conditions without SiO 2 , occupying approximately 7.9% of the total Cl-components at temperature of 950 o C. Besides, thermodynamic simulation confirms the formation of the metal-matrix compounds, i.e. Na 2 Si 2 O 5 (s) as shown in Figure 4.3 (d), and its molar fraction in

  It can be concluded that the overall conversion rate of chlorine from NaCl to Cl -ion is less than 8% even at higher temperature of 950 °C and with the presence of water and SiO 2 . The amount of O 2 content slightly increases the conversion from NaCl to Cl -ion; therefore, the effect of Chapter 4: Influence of CaO reactant on HCl release and control from inorganic and organic sources of chlorine 98To sum up, results obtained from tubular furnace experiments are in good accordance with that from thermodynamic equilibrium simulation, indicating that the presence of O 2

	atmosphere on chlorine conversion rate from NaCl in descending order is: O 2 > air > N 2
	(conversion rates at 950 °C are 3.2% > 2.7% > 1.2% respectively). The presence of water and
	SiO

2 significantly increases the chlorine conversion from NaCl to Cl -ion. At temperature of 950 °C, the conversion rate is increased by 1.6% and 5.1% for the cases with the presence of water alone and with both water and SiO 2 .

Table 4 . 1 Characteristic parameters of PVC pyrolysis

 4

	Material

  34.3% -61.3% for Ca/Cl ratio of 1:2 and 51.2% -82.4% for Ca/Cl ratio of 1:1 under experimental temperature. Therefore, excessive use of CaO reactant than stoichiometric amount is essential to achieve acceptable HCl mitigation efficiency levels.Chapter 4: Influence of CaO reactant on HCl release and control from inorganic and organic sources of chlorine 104 compared to that of the Ca/Cl molar ratio of 1:2. However, when the ratio is further increased to 4:1, a limited enhancement of HCl mitigation efficiency by around 5% is observed. The reason may be attributed to the fact that at higher Ca/Cl molar ratio conditions, there are enough active sites on CaO reactant particle surface for dechlorination reactions. This means that under such circumstance, the amount of CaO (or in other words, the number of the active sites on CaO surface) is no longer the limiting factor of the reactions; therefore, further increasing Ca/Cl molar ratio will lead to a slight improvement on HCl mitigation efficiency. Accordingly, the Ca/Cl molar ratio of 2:1 is believed the most effective amount of CaO for HCl gas removal under the experimental conditions. Fontana et al.[16]and Tan et al.[17] also found that the increase of the Ca/Cl molar ratio will enhance HCl mitigation effect of Ca-based reactants and the most proper Ca/Cl molar ratio is in the range of 2:1 to 3:1.

	CaO HCl 		CaClOH	(4.7)
	2 CaO HCl CaCl H O 2 2   	(4.8)

As exhibited in

Figure 4.10, when the Ca/Cl molar ratio reaches 2:1, HCl mitigation efficiency ranges from 92.4% at temperature of 550 o C to 72.2% at temperature of 850 o C, resulting in a remarkable enhancement of HCl mitigation efficiency by around 31% to 38% as With respect to CaO conversion rate, the influences of Ca/Cl molar ratio on CaO conversion rate during PVC pyrolysis are also reported in Figure 4. 10. As illustrated, although

HCl mitigation efficiency is improved with increasing Ca/Cl molar ratio, the corresponding CaO conversion rate declines continuously at the same time. At the Ca/Cl molar ratio of 1:2, the highest CaO conversion rate of 34.3% -61.3% is obtained; however, at the Ca/Cl molar ratio of 4:1, CaO conversion rate declines significantly to 9.8% -11.9%. The reduction of CaO conversion rate is reasonable, since with the increase of Ca/Cl molar ratio, the amount of CaO is more abundant. Taking into consideration at the working condition of the most effective Ca/Cl molar ratio of 2:1, CaO conversion rate is as low as 18.1% -23.1%, the reuse of the spent in-furnace CaO reactant together with the fresh CaO can be regarded as a promising alternative to reduce the sorbent consumption.

Table 4 . 2 Elemental composition of the spent CaO reactant by EDS analysis

 4

	Elements	Carbon (C) a	Oxygen (O) a	Chlorine (Cl) a	Calcium (Ca) a	Total
	Ca/Cl=1:1 b	34.0	22. 2	17.6	26.2	100
	Ca/Cl=2:1 b	23.7	32.3	8.6	35.4	100
	a in atomic percentage;
	b at temperature of 750 o C.
	As presented in Figure 4. 11 and Table 4.2, at a temperature of 750 o C, chlorine content
	in the residues obtained from the cases of Ca/Cl molar ratio at 1:1 and 2:1 is 17.6% and 8.6%,
	respectively, which further confirms that although HCl mitigation efficiency is improved at
	higher Ca/Cl molar ratio, the conversion of CaO to CaCl 2 (reflected by chlorine content in the
	residues) is decreased.

Table 4 . 3 Elemental composition of the spent CaO reactant by EDS analysis

 4

	Elements	Carbon (C) a	Oxygen (O) a	Chlorine (Cl) a	Calcium (Ca) a
	550 o C b	40.5	21.2	20.4	17.9
	650 o C b	30.1	25.3	12.6	26.0
	750 o C b	23.7	32.3	8.6	35.4
	850 o C b	22.2	35.3	4.0	38.5
	a in atomic percentage;

b at Ca/Cl molar ratio of 2:1.

  14, when 80-120 mesh CaO reactant with the Ca/Cl molar ratio of 1:2 is used, HCl mitigation efficiency ranges from 52.3% to 84.3%, nearly reaching the same HCl mitigation efficiency level as that when 60-80 mesh CaO reactant with the Ca/Cl molar ratio of 1:1 is used (51.2% -82.4%). Similarly, when 80-120 mesh CaO reactant with the Ca/Cl molar ratio of 1:1 is used, HCl mitigation efficiency is as high as 66.35% -98.2% under the experimental temperature.These efficiencies are much closed to the aforementioned most preferable HCl mitigation conditions using the CaO reactant with the particle size of 60-80 mesh and the Ca/Cl molar

ratio of 2:1 in Section 4.3.2 (72.2%-92.4%).

Table 4 . 4 Fraser-Suzuki deconvolution results for PVC pyrolysis with/without CaO reactant at 5, 10 and 20 °C/min

 4

	Single			PVC		Single		PVC with CaO
	reaction		5	10	20	reaction		5	10	20
		H p	-2.25	-2.00	-1.95		H p	-2.00	-1.75	-1.65
		T p	272.0 0	286.2 0	297.9 0		T p	276.8 0	287.9 0	299.2 0
	FSF 1 :	W h				F"F 1 :	W h
	de-HCl	f	22.00 24.00 24.00	de-HCl	f	18.00 20.00 20.00
		A s	-0.10	-0.20	-0.30		A s	-0.20	-0.30	-0.30
		c i	52.88 51.83 51.46		c i	38.88 38.48 36.27
		H p	-0.41	-0.29	-0.30		H p	-0.30	-0.29	-0.30
	FSF 2 : de-benzen e	T p W h f	314.9 0 38.00 45.00 48.00 325.3 0 336.0 0	F"F 2 : de-benzen e	T p W h f	306.9 0 50.00 50.00 50.00 320.0 329.5 0 0
		A s	-0.41	0.30	0.30		A s	0.30	0.30	0.40
		c i	16.83 14.35 11.83		c i	16.49 15.94 16.92
		H p	-0.48	-0.40	-0.38		H p	-0.75	-0.71	-0.65
	FSF 3 :	T p	446.6 0	459.5 0	469.0 0	F"F 3 :	T p	442.1 0	452.8 0	466.4 0
	de-polyene	W h				de-polyene	W h
	and	f	65.00 65.00 80.00	and	f	55.00 60.00 65.00
	cyclization	A s	-0.10	0.01	-0.10	cyclization	A s	-0.05	0.10	-0.05
		c i	33.33 27.68 32.48		c i	43.95 45.51 45.01
	DEV (%)		2.99	1.84	2.65			1.90	2.37	2.36
	4.4.2.

2 Activation energy estimation

  l β/T 2 ) vs 1/T curves for KAS method at the same conversion rate of 0.1, 0.2,

	0.3, 0.4, 0.5, 0.6, 0.7, 0.8, and 0.9.
	(a) FSF 1 of OFW method	(b) FSF 1 of KAS method
	(c) FSF 2 of OFW method	(d) FSF 2 of KAS method
	(e) FSF 3 of OFW method	(f) FSF 3 of KAS method

Following the separation of the DTG peaks by Fraser-Suzuki deconvolution, the conversion rate of each pseudo-component is determined. And then according to Section 4.4.1.4, the ki eti pa a ete s a e o tai ed plotti g l β s /T u es fo OFW method a d

Table 4 . 5 The characteristic temperature at each conversion rate of different pseudo-component, the slope of the regression plots and the estimated activation energy of PVC pyrolysis without/with CaO reactant PVC FSF 1

 4

			Characteristic temperature, o C		OFW method		KAS method
	Conversion rate	5	10	20	Slope	Ea	Dev% a	Slope	Ea	Dev% a
	a = 0.1	259.00 269.34 278.85 -20.49 170.38 11.46 -19.41 153.39 12.36
	a = 0.2	263.27 274.85 285.14 -18.94 157.47 3.01 -17.85 141.04 3.31
	a = 0.3	266.63 278.72 289.09 -18.68 155.28 1.58 -17.58 138.90 1.74
	a = 0.4	269.04 282.01 292.47 -18.05 150.04 -1.84 -16.94 133.88 -1.94
	a = 0.5	271.41 284.76 295.28 -17.86 148.47 -2.87 -16.75 132.34 -3.06
	a = 0.6	273.81 287.52 298.08 -17.71 147.24 -3.68 -16.59 131.14 -3.95
	a = 0.7	276.24 289.74 300.33 -18.01 149.74 -2.04 -16.89 133.47 -2.23
	a = 0.8	279.16 293.16 303.74 -17.82 148.16 -3.07 -16.69 131.93 -3.37
	a = 0.9	282.56 297.07 307.20 -17.92 148.95 -2.56 -16.78 132.61 -2.86
	E a,mean					152.86			136.52
	PVC FSF 2		Characteristic temperature, o C		OFW method		KAS method
	Conversion rate	5	10	20	Slope	Ea	Dev% a	Slope	Ea	Dev% a
	a = 0.1	301.50 308.36 317.58 -29.08 241.77 13.11 -27.91 220.61 13.90
	a = 0.2	307.32 315.08 325.08 -26.91 223.74 4.67 -25.81 203.99 5.33
	a = 0.3	312.11 320.68 330.82 -26.15 217.44 1.73 -24.97 197.30 1.87
	a = 0.4	316.42 325.14 335.94 -25.43 211.42 -1.09 -24.23 191.49 -1.13
	a = 0.5	321.23 330.69 341.65 -24.79 206.11 -3.57 -23.58 186.37 -3.77
	a = 0.6	326.50 335.69 347.33 -24.68 205.16 -4.02 -23.46 185.38 -4.28
	a = 0.7	332.31 341.76 353.58 -24.66 205.05 -4.07 -23.43 185.18 -4.39

Figure 4. 17 and Figure 4. 18 present the kinetic plots of the pseudo-components from PVC pyrolysis without/with CaO reactant. Correlation coefficients of the linear regression, R 2 , are higher than 0.90, and the plotted trend lines at different conversion rates are almost parallel, indicating that the calculation results are reliable [30]. The characteristic temperature at each conversion rate of different pseudo-component, the slope of the regression plots and the estimated activation energy are listed in Table 4.5. Accordingly, the De % alues of the o tai ed Eα a e less tha

. % at all the conversion rates of each pseudo-components, which means, according to ICTAC Kinetics Committee recommendations, the reaction is mainly governed by a single reaction mechanism and can thus be adequately described by a single-step model

[20, 33]

.

As listed in Table

4

.5, the average apparent activation energies of the considered pseudo-components for PVC pyrolysis are: FSF 1 (de-HCl), 152.86 and 136.52 kJ/mol; FSF 2 (de-benzene), 213.75 and 193.68 kJ/mol, and FSF 3 (de-polyene and cyclization), 308.36 and 281.59 kJ/mol according to OFW and KAS methods, respectively. The obtained results are pyrolysis. The results are in good agreement with the apparent activation energy estimation results; because the average Ea is increased by approximately 12% when CaO is mixed with

  Cl molar ratio or reducing the particle size significantly enhances HCl mitigation efficiency. At temperature of 750 o C, HCl mitigation efficiency reaches 85.6% and 83.35% when 40-60 mesh CaO with Ca/Cl molar ratio of 2:1 and 80-120 mesh CaO with Ca/Cl molar ratio of 1:1 are used.

	 SEM and EDS analysis confirm that HCl mitigation by CaO reactant is limited by
	both chemical reaction rate and physical surface properties. At higher temperature,
	the chemical reaction equilibrium shifts to the back forward direction of
	dechlorination reactions; meanwhile, calcination and sintering of the reactants
	significantly reduce the specific surface area and porosity of the reactant, resulting
	in the reduction of HCl mitigation efficiency at higher temperature.
	 HCl release from organic source PVC onsets at low temperature of around 200 o C
	and ends up at around 400 o C. HCl is the main products of the first degradation
	stage for PVC pyrolysis and nearly all the chlorine contained in PVC is discharged as
	HCl via dehydrochlorination reaction in this stage.

Both thermodynamic equilibrium simulation and experimental study confirm that less than 1.2% of the total chlorine is converted from NaCl to Cl-content species during pyrolysis within the considered temperature range. The presence of O 2 , water, and inorganic materials like SiO 2 are able to promote the conversion of NaCl to Cl-content contaminants. However, under the experimental conditions, the overall conversion rate of Cl-contaminants is less than 8%.  In-furnace CaO reactant is effective for HCl mitigation during PVC pyrolysis. Higher temperature plays adverse effects on HCl mitigation efficiency. Increasing the Ca/ The estimated average apparent activation energy of the three pseudocomponents for PVC pyrolysis is: FSF 1 (de-HCl), 152.86 and 136.52 kJ/mol; FSF 2 (de-benzene), 213.75 and 193.68 kJ/mol, and FSF 3 (de-polyene and cyclization), 308.36 and 281.59 kJ/mol, by using OFW and KAS calculation method, respectively. The use of CaO reactant increases the apparent activation energy of dehydrochlorination reaction but decreases the average Ea of hydrocarbon decomposition reactions. The use of CaO decreased the apparent activation energy of the overall PVC decomposition reaction from 197.26 to 148.92 kJ/mol.

Table 5 . 1 Identified tar compositions and their relative contents from Py-GC/MS experiments of food waste at 750 o C with or without CaO reactant

 5

	Time	Pyrolytic tar from food waste without CaO	Pyrolytic tar from food waste with CaO
	(min)	Formula	Species	Area% Formula	Species	Area %
	2.05	*	*	*	C5H6	1,3-Cyclopentadie ne	7.30
	2.28	C5H11N Pent-4-enylamine	3.81	*	*	*
	2.35	C5H8	1,4-Pentadiene	2.43	*	*	*
	2.42	C2H4O Acetaldehyde	9.83	*	*	*
	2.58	C5H6	1,3-Cyclopentadiene	4.91	C3H6O	Trimethylene oxide	2.03
	2.81	*	*	*	C6H8O2 2-Hexynoic acid	1.41
	2.85	C4H4O Furan	5.84	C3H6O Acetone	8.59
	2.99	C3H6O Acetone	6.29	C6H8	1,3-Cyclopentadie ne, 5-methyl	1.76
	3.08	*	*	*	C6H8	1,3-Cyclopentadie ne, 5-methyl	2.01
	3.17	*	*	*	C3H4O 2-Propenal	1.88
	3.23	C3H4O 2-Propenal	3.61	C4H8O 2-Butanone	4.80
	3.45	C5H6O Furan, 3-methyl	2.99	*	*	*
	3.55	C4H6O Methacrolein	2.39	*	*	*
	3.77	C4H8O 2-Butanone	1.81	C6H6	Benzene	5.41
	4.30	C6H6	Benzene	3.74	C4H6O	Methyl vinyl ketone	1.89
	4.37	C4H6O Methyl vinyl ketone	3.65	C5H10O 2-Pentanone	3.26
	4.47	C6H8O Furan, 2,5-dimethyl	1.76	*	*	*
	4.82	C4H6O2 2,3-Butanedione	2.41	C3H3N 2-Propenenitrile	1.56
	5.06	*	*	*	C2H3N Acetonitrile	3.87
	5.27	C7H8	Toluene	2.43	C7H8	Toluene	8.41
	5.98	C2H3N Acetonitrile	7.48	C8H10	Ethylbenzene	2.98
	7.86	C8H10 Ethylbenzene	2.34	C8H10	o-Xylene	1.60
	8.16	*	*	*	C5H8O Cyclopentanone	1.87
	9.31	*	*	*	C8H8	Styrene	2.04
	11.11	*	*	*	C3H6O2	2-Propanone, 1-hydroxy	3.60
	12.27 C3H6O2	2-Propanone, 1-hydroxy	4.14	C5H6O	2-Cyclopenten-1-one	3.12
	13.74	C5H6O 2-Cyclopenten-1-one	2.71	C6H8O	2-Cyclopenten-1-one, 2-methyl	2.00
	14.09	*	*	*	C2H4O2 Acetic acid	2.33
	16.10 C2H4O2 Acetic acid	3.78	C5H4O2 Furfural	1.80
	16.42 C5H4O2 3-Furaldehyde	3.97	C9H8	Indene	1.53

To better reflect the effect of CaO reactant on pyrolytic tar compositions, the detected components are classified into 10 groups as olefins, nitriles, amines, benzenes, furans, aldehydes, ketones, phenols, acid compounds, and others (refers to nitrogen-content heterocyclics and oxides in this section). The distributions of the detected species are exhibited in Figure

5

.7. As presented, oxygenated compounds are the dominant species contained in pyrolytic tar, accounting to 70.5% and 55.1% under non-catalytic and catalytic conditions, followed by aromatic compounds (8.5% and 23.9%, respectively) and hydrocarbons (7.3% and 11.1%, respectively). The role of CaO reactant is obvious on oxygen-content compounds elimination and aromatic compounds promotion. By adding CaO reactant, the relative content of oxygenated compounds is decreased by 21.9%; in contrast,

Table 5 . 2 Identified tar compositions and their relative contents from Py-GC/MS experiments of PVC at 750 o C with or without CaO reactant

 5

	Pyrolytic tar compounds from PVC without CaO	Pyrolytic tar compounds from PVC withCaO
	Time (min)	Formula	Species	Area %	Time (min)	Formula	Species	Area %
	1.97	C3H8	Propane	3.02	1.84	CO2	Carbon dioxide	2.45
	2.04	C4H6	Methylene-cyclopropane	1.52	1.97	C4H6	Methylene-cyclopropane	1.06
	2.24	C5H8	1,3-Pentadiene	0.84	2.11 C6H8O2	2-Hexynoic acid	0.99
	2.45	C5H6	Cyclopropane, ethynyl-	1.94	2.32	C5H6	1,3-Cyclopenta diene	1.61
							1,3-Cyclopenta
	4.07	C6H6	Benzene	16.00	3.08	C6H8	diene,	0.65
							5-methyl-
	4.26	HCl	Hydrogen chloride	3.18	3.96	C6H6	Benzene	14.64
	5.68	C7H8	Toluene	9.81	5.6	C7H8	Toluene	8.13
	7.52	C8H10 Ethylbenzene	1.23	7.44	C8H10 Ethylbenzene	1.01
	7.7	C8H10 o-Xylene	0.71	7.78	C8H10 o-Xylene	0.61
	7.85	C8H10 p-Xylene	1.07	8.85	C8H10 o-Xylene	1.28
	8.9	C8H10 o-Xylene	1.87	10.7 1	C8H8	Bicyclo[4.2.0]o -triene cta-1,3,5	4.00
	10.74	C8H8	Bicyclo[4.2.0] octa-1,3,5-trien e	3.91	10.9 2	C9H10	Benzene, cyclopropyl-	0.66
	13.35	C9H10	Benzene, 1-ethenyl-3-met	1.09	13.3 3	C9H10	Benzene, cyclopropyl-	0.89

Table 6 .

 6 1 lists the experimental results of the effect of temperature and CaO reactant on product mass distribution. The results are presented as mass ratio of the specific fraction (gas, tar, or solid residues) to the total mass of the products (i.e. the sum of gas, tar, and solid residues). Results reveal that temperature significantly influences the mass distribution of According to Table6.1, when pyrolysis temperature is increased from 550 to 850 o C, yields decline significantly from 46.8 to 18.6 wt.% for tar and from 20.4 to 12.6 wt.% for solid residues; simultaneously, produced gas yield is increased continuously from 32.8 to 68.6 wt.%. The results agree well with literature data[20][21][22]. And it is quite reasonable, since temperature plays an important role in chemical reactions by shifting the thermodynamic equilibrium. With special regards to waste pyrolysis in the present work, increasing the temperature favors the emitting of volatiles from S-MSW at initial waste decomposition stage. And meanwhile, high temperature also significantly facilitates tar thermal cracking reactions, resulting in the reduction of tar yields and the production of more incondensable gases.

gas, tar and solid residues from S-MSW pyrolysis. Higher temperature improved the mass percentage of produced gas but decreased simultaneously the proportions of tar and solid Chapter 6: Influence of CaO reactant on pyrolysis of simulated MSW: HCl mitigation and tar upgrading 157 residues.

Table 6 . 1 Effect of temperature and CaO reactant on pyrolysis product mass distribution

 6

	Temperature, o C	Reactant	Solid products, wt. % a Tar, wt. % Gas products, wt. %
	550	Without CaO With CaO	20.4 25.4	46.8 37.8	32.8 36.8
	650	Without CaO With CaO	15.8 21.6	35.0 24.0	49.2 54.4
	750	Without CaO With CaO	15.0 18.6	30.6 15.2	54.4 66.2
	850	Without CaO With CaO	12.6 17.6	18.6 14.0	68.8 68.4
	promotes tar catalytic thermal cracking reactions and facilitates the conversion from bio-oil
	to light gas products. The yield of produced gas is also increased, except for experiments at
	temperature of 850 o C, mainly due to the enhancement of S-MSW decomposition and tar
	cracking with the addition of CaO reactant. With respect to solid fractions, the generation of
	solid residues is increased by 3.6 % -5.8 % under the considered temperature ranges;

a the mass of raw CaO added is subtracted from the solid residues.

When CaO reactant is co-fed with S-MSW, tar yield is significantly reduced from 4.6 to 15.4 wt. % under the experimental temperature ranges. The results are in good accordance with the findings presented in Chapter 5, revealing that the presence of CaO reactant

Table 6 . 2 Reactions related to MSW pyrolysis and CaO reactant

 6

	Reactions related to MSW pyrolysis
	General pyrolysis	m n CH O			H	2				2 CO CO  			a C H tar residues b  	6.1
	Water-gas (i)	mn CH O		  1-n H O CO 2 			-2 2 2 mn 		H	2	6.2
	Water-gas (ii)	2 C H O CO H   	2	6.3
	Water-gas shift	2 CO H O CO H 2   	2	6.4
	Hydrocarbon reforming	2 C H aH O ab 		aCO		2	2 ab 	H	2	6.5
	Tar thermal cracking	Tar	Heat   		H	2			2 CO CO hydrocarbons   	 	6.6
	Reactions related to CaO reactant
	Carbonation/calcination	2 CaO CO 			3 CaCO	6.7
	Dechlorination (i)	CaO HCl 		CaClOH	6.8
	Dechlorination (ii)	2 CaO HCl 		2 CaCl H O 2 	6.9
	Dechlorination (iii)	3 CaCO HCl 			2 CaClOH CO 	6.10
	Dechlorination (vi)	3 CaCO		2	HCl			2 CaCl H O CO 2 2  	6.11
	Tar catalytic thermal cracking	Tar	CaO   	H	2			2 CO CO hydrocarbons   	 	6.12

Table 6 . 3 Typical tar compounds detected by GC/MS derived from S-MSW pyrolysis at 550 o C with or without CaO reactant

 6

	Time	550 o C pyrolysis without CaO	550 o C pyrolysis witht CaO
	(min)	Formula	Species	Area % Formula	Species	Area %
			Cyclopentane,
	5.80	C10H18	1-ethenyl-3-eth	1.18	*	*	*
			yl-2-methyl
	6.31	C8H16O	3-Octen-1-ol	1.32	C11H20	1,6-Undecadiene	2.32
	6.55	C10H20	1-Decene	7.08	C10H20 1-Decene	8.07
	6.78	C10H22	Decane	0.93	*	*	*
	7.75	C10H16O	cis-p-mentha-1( 7), S-dien-2-ol	0.98	*	*	*
	8.28	C10H16O	cis-p-mentha-1( 7), S-dien-2-ol	1.25	*	*	*
	8.69	C14H28O	E-7-Tetradecenol	2.30	*	*	*
	8.70	*	*	*	C13H24	1,12-Tridecadiene	2.46
	8.87	C11H22	1-Undecene	6.15	C11H22 1-Undecene	6.32
	9.04	C11H22O	trans-2-Undecen-1-ol	4.71	*	*	*
			Cyclohexanone,
	9.31	C11H18O2	2,2-dimethyl-5 -(3-methyloxiran	3.00	*	*	*
			y)-
	10.10	*	*	*	C10H10	1H-Indene, 3-methyl-	2.95
	10.44	C10H8	Naphthalene	2.94	C10H8	Naphthalene	4.23
	10.52	C10H12O4 Cantharidin	1.18	*	*	*
	10.66	C12H24	3-Dodecene	5.05	C12H24 3-Dodecene	5.60
						Hydrocinnamic
						acid,
	11.76	*	*	*	C20H22O 2	o-[1,2,3,4-tetr ahydro-2-naph	4.39
						thy)methyl)me
						thyl]-
	12.08	*	*	*	C13H24	1,12-Tridecadiene	2.19
	12.20	C13H26	2-Trifluoroaceto xy tridecane	4.31	C14H28 7-Tetradecene	6.20
	12.43	*	*	*	C11H10	2-methyl-Naphthalene	2.91
	12.76	*	*	*	C20H22O Hydrocinnamic	2.28

Table 6 . 4 Typical tar compounds detected by GC/MS derived from S-MSW pyrolysis at 750 o C with or without CaO reactant

 6

	Time	750 o C pyrolysis without CaO	750 o C pyrolysis with CaO
	(min)	Formula	Species	Area % Formula	Species	Area %
			Benzene,
	6.57	C9H10	1-ethenyl-4	1.89	*	*	*
			-methyl-
						Benzene,
	7.74	C9H8	Indene	3.08	C9H8	1-ethenyl-4	5.85
						-methyl-
	9.98	C10H10	Naphthalene, 1,2-dihydro-	3.01	C10H10	Naphthalene, 1,2-dihydro-	3.48
	10.44	C10H8 Naphthalene	27.19	C10H8	Naphthalene	31.13
	12.19	C11H10	Naphthalene, 2-methyl-	8.15	C11H10	Naphthalene, 2-methyl-	8.36
	12.43	C11H10	Naphthalene, 2-methyl-	6.84	C11H10	Naphthalene, 2-methyl-	6.65
	13.40	C12H10 Biphenyl	4.36	C12H10 Biphenyl	2.75
	13.58	C12H12	Naphthalene, 1,3-dimethyl-	1.60	*	*	*
	13.92	C12H12	Naphthalene, 2,6-dimethyl-	2.30	*	*	*
	14.07	C12H12	Benzene, 2,5-cyclohexadien-1-yl-	4.12	C12H10	Naphthalene, 2-ethenyl-	3.89
	14.30	C12H8 Acenaphthylene	6.21	C12H8	Acenaphthylene	5.66
	14.77	C13H12	1-Isopropeny1 naphthalene	2.15	C13H12	1-Isopropeny1 naphthalene	2.16
	15.94	C13H10 Fluorene	5.39	C13H10 Fluorene	5.18
	16.18	*	*	*	C13H12	Fluorene,2,4a -dihydro-	2.10
	18.14	C14H10 Anthracene	7.18	C14H10 Phenanthrene	8.51
	18.23	C14H10 Anthracene	2.98	C14H10 Phenanthrene	3.66
	19.55	C15H12	Phenanthrene, 1-methyl-	2.84	C15H12	Phenanthrene, 1-methyl-	3.27
						Pyrene,
	20.68	*	*	*	C16H14	4,5,9,10-	1.88
						-tetrahydro-
	21.12 C19H21N Norrtriptyline	2.90	C19H21N Norrtriptyline	1.34
	21.37	C16H10 Fluoranthene	2.69	C16H10 Fluoranthene	2.08
	22.26	C17H12	11H-Benzo [b]fluorene	2.55	C17H12	11H-Benzo [b]fluorene	2.04
	22.42	C17H12	Fluoranthene, 2-methyl-	2.56	*	*	*

Table 7 . 1 Waste gasification plants running in Europe, data source from Molino et al.

 7

	[18]

Table 7 . 2 Physical compositions of MSW

 7

		MSW composition (wt. %)
		Bridgeable waste Paper Plastic Textile Glass Metal Other a
	European average	35.0	22.0	10.0	3.0	6.0	4.0	20.0

a othe aste i lude i ed te tiles, leathe , u e , laminates, e-waste, inert materials (e.g. gravel, ceramics, ash), etc.

Table 7 . 3 Material and energy flows for the considered systems (data are presented based on 1 t/MSW)

 7

	Unit	S1	S2	S3	S4

Table 7 .

 7

		4 The E.U. average electricity matrix (Unit: %)
		EU average electricity matrix, %
	Fossil fuels	47.2
	Nuclear	24.09
	Hydroelectric	16.64
	Geothermal	0.33
	Solar	1.99
	Tide and wave	0.01
	Wind	5.92
	Biomass and waste	4.14

  except for S4; savings of AP from etals e li g of S4 is one order of magnitude lower than that of other systems. The reason is mainly attributed to the fact that aluminum coming from gasification-melting process is present as amorphous phase and is

	As o pa ed to " o e tio al i i e atio , GWP e issio f o	ate ials a d
	e e g p o isio of p o-gasification systems (S1, S2, and S3) is 2-3 times higher. It is mainly
	due to the additional energy and materials consumption in these systems, for example, MSW
	pre-treatment of S3, and limestone and coke addition of S4. As aforementioned, energy
	recovery a d	etal e li g a e the e i o e tal e efits f o WtE s ste s. " leads
	to the highest e i o e tal sa i gs of GWP f o e e g e o e		a o g the o side ed
	systems, mainly attributed to the highest net electricity generation efficiency of 27.4% as
	aforementioned. E issio s f o	otto ash la dfill of " a d " , slag e li g of S2
	a d " , APC esidues sta ilizatio a d la dfill , a d t a spo tatio	ause egligi le GWP
	impacts of all the considered systems, ranging from -0.6 to 5.0 kg CO 2 eq.
	I ega d to the i pa t atego of AP, sta k e issio is also fou d the ajo
	co t i uto of AP e issio . Ho e e , AP e issio f o sta k e issio of " is u h lo e
	than that of S1, S2, and S3, mainly attributed to the reduction of NO x emission from
	gasification-melting-based WtE system as presented in Table 7.3. Comparable environmental
	sa i gs of AP f o	etals e li g is o se ed as that f o	e e g e o e	of the
	considered systems,

Table 7 . 6 Material and energy flows for the considered systems (data are presented based on 1 t/MSW)

 7

	Unit	"	S3	S4

Table 7 . 7 Characterized environmental impacts of the considered system

 7

				S4:
		S3': Hypothetical	S3:
	Categories and Unit			Gasification-
		gasification	Gasification
				melting
	GWP [kg CO2 eq.]	56.12	104.00	422.00
	Non-toxic
	AP [kg SO2 eq.]	-0.64	-0.51	-0.43
	impacts
	EP [kg Phosphate eq.]	0.03	0.04	-2.69E-03
	TETP [kg DCB eq.]	-0.24	-0.23	2.65
	Toxic
	FAETP [kg DCB eq.]	-5.00	-4.94	-7.77
	impacts
	HTP [kg DCB eq.]	-203.89	-202.00	-11.50
	a the best performance of the impact categories are represented by bold type and
	marked as red.

0.1 Philosophie du travail de recherche : étudier le captage et l'i o ilisatio i -situ du chlore émis lors de la pyro-gazéification de résidus organiques solides en utilisant un réactif CaO La

  Ainsi, ces résidus solides contiennent des proportions non négligeables de sources chlorées, de atu e o ga i ue plasti ue ou i o ga i ue sel . Lo s d u t aite e t the o hi i ue de o pośs hlo ́s, de l acide chlorhydrique (HCl) est formé, conduisant à des désagréments tels ue la o osio , la fo atio de o ta i a ts to i ues, l a idifi atio … La Figu e . . Ǵ ́ e u s gas de ualit́ suṕ ieu e, e o sid́ a t u u e pa tie i po ta te de polluants a été éliminée, et que le ratio H 2 /CO peut aussi être amélioré. En effet, le but de la pyro-gazéification est de fournir un gaz énergétique appelé syngas, composé majoritairement de H 2 et de CO. Les proportions de chacun de ces 2 gaz, conduisant à des désagréments tels que des pluies acides formées localement, la corrosion à haute te ṕ atu e des haudiè es e ui ́duit de a iè e sig ifi ati e l effi a it́ à la o e sio ́ e ǵti ue des i stallatio s d i i ́ atio , la fo atio de o ta i a ts toxiques tels que les dibenzo-p-dioxines et les dibenzofuranes polychlorés. De ce fait, il o ie t o t ôle l ́ issio de HCl da s es p o ́d́s afi de li ite les is ues de o ta i atio pa des ́ĺ e ts hlo ́s, et ie sû espe te les li ites d ́ issio imposées par les réglementations en vigueur. Actuellement, les outils de contrôle des émissions de HCl existants sont : (1) le contrôle de la quantité de chlore dans les résidus solide avant traitement thermochimique, en utilisant par exemple le lavage, le tri, des p ́t aite e ts à asse te ṕ atu e… ; (2) le contrôle de la quantité de chlore sous forme d HCl da s le gaz ́ is ap ès le t aite e t the o hi i ue t aite e t pa oie sè he ou voie humide, ou un procédé intermédiaire) ; (3) le contrôle peut être réalisé au cours du t aite e t the o hi i ue. C est p ́ iśment cette dernière option de contrôle in-situ qui a été sélectionnée pour la présente étude. te, l ́ aluatio de l i pa t e i o e e tal e ǵ ́ al des pa a èt es tels ue l a idifi atio , l ́ oto i it́, l i pa t su le ha ge e t li ati ue … so t ́ alús et e fi la de iè e ́tape o stitue l i te p ́tatio des ́sultats.

	-Eviter/limiter les phénomènes de corrosion des installations initiés par la présence
	du gaz HCl émis.
	-Limiter la formation de dioxines, qui résultent de réactions hétérogènes entre le
	résidu solide et la p ́se e de hlo e à l ́tat gazeu	-400°C). Les dioxines sont
	des pollua ts to i ues u il o ie t d ́li i e pou u e utilisatio ult́ ieu e du
	syngas.
	Résumé du Chapitre 1 Etat de l'a t -La se tio . i t oduit l effet de l utilisatio de CaO o	e ́a tif i -situ sur
	p ́se te la philosophie ǵ ́ ale du t a ail, à sa oi l utilisatio i -situ du réactif CaO lors de Ainsi, le présent travail de thèse est focalisé sur « l a atte e t des o ta i a ts
	la pyro-gazéification des résidus solides. La présence de ce réactif in-situ a pour objectifs de : chlorés lors de la pyro-gazéification de déchets en utilisant un réactif de CaO : étude
	-Mod́ e l ́ issio de HCl lo s du p o ́d́ de p o-gazéification des résidus solides expérimentale et analyse du cycle de vie ». La structure du manuscrit est présentée à la
	figure 1.0.2.

L o je tif de e hapit e est de p ́se te l ́tat de l a t de la recherche bibliographique et les motivations de ce travail de recherche. Ainsi, sont présentés et discutés les progrès des recherches en cours concernant : l ́ issio de HCl lo s de la o e sio the o hi i ue des déchets municipaux et les méthodes de contrôle associées ; le rôle joué par le réactif CaO sur le craquage et la valorisation des goudrons lors de son utilisation in-situ dans le réacteur ; et la ́thodologie d a al se de le de ie ACV et so appli atio au s stè es de conversion thermochimique des déchets municipaux.

Ainsi, la section 1.1 détaille la motivation du présent t a ail ui est l a ́lio atio si ulta ́e de l effi a it́ ́ e ǵti ue et de la pe fo a e e i o e e tale des p o ́d́s de traitement thermique de ces résidus solides. Les déchets municipaux sont constitués, en fractions variables, de déchets organiques, papier, plastiques, verre, métaux et textiles.

ui se d́ oule à hautes te ṕ atu es. Le ut est d i o ilise le hlore dans la f a tio solide esta t à l i t́ ieu du ́a teu . mais le taux et la nature des polluants, orientent à la fois son utilisation subséquente et les méthodes de purification nécessaires.

Figure 1. se tio . p ́se te les a a t́ isti ues de l ́ issio de HCl et les ́thodes de contrôle appliquées aux technologies de traitement thermique des déchets municipaux. La quantité de chlore présente dans ces résidus est relativement élevée (de 800 à 2500 mg/kg). Ai si le t aite e t the o hi i ue o duit à la fo atio d a ide hlo h d i ue HCl , l a ́lio atio de la qualité du syngas produit lors de la pyro-gazéification, et sur le craquage des goudrons. En résumé, CaO permet de capter partiellement le CO 2 , un gaz non ́ e ǵti ue, et ai si d a ́lio e la ualit́ ́ e ǵti ue du s gas. De plus CaO pa ti ipe au a uage des goud o s, pe etta t d aug e te le tau de o e sion de la ressource en syngas. La se tio . d́ it l app o he et la ́thodologie glo ale d u e a al se de le de ie et sa ise e oeu e da s le ad e du t aite e t the i ue des d́ hets u i ipau . Le travail est découpé en 4 phases : la définition des o je tifs et du ha p d i estigatio , l i e tai e des do ́es et des flu asso ís au f o tiè es d́fi ies lo s de la phase p ́ ́de

Figure 1.

0.2 Structure et principaux aspects traités dans la thèse

		Etat de l a t
		Chapitre 1
		Matériels et méthodes
		Chapitre 2
	Mécanisme de	Emission de	Rôle du réactif
	réaction	HCl par des	CaO sur le
	gaz/solide	sources	craquage
	entre CaO et	organiques et	pyrolytique et
	HCl	inorganiques	la valorisation
	étude	et utilisation	des goudrons,
	expérimentale	du réactif CaO	en utilisant
	et modélisation cinétique	in-situ pour des émissions l a atte e t	des composés modèles
		de HCl
	Chapitre 3	Chapitre 4

Chapitre 5 Etude expérimentale et modélisation cinétique

Influence du réactif CaO in-situ lors de la pyrolyse des déchets municipaux : abattement de HCl et craquage des goudrons Chapitre 6

Analyse de cycle de vie des technologies de Waste-to-Energy : systèmes de pyro-gazéification et prise en compte de différentes optio s d a atte e t des ́ issio s du hlo e

Chapitre 7 Conclusions et perspectives Chapitre 8 Modèle d' valuatio Résumé du Chapitre 2 Matériels et méthodes

  Dans la section 1.2, les caractéristiques des déchets municipaux considérés et du réactif CaO sont introduites. Ce sont 5 composés modèles, représentatifs des 5 grandes classes de d́ hets u i ipau , ui so t śle tio ́s et a a t́ iśs pou la ise e oeu e de l ́tude : du PVC et du PE (pour représenter les plastiques et en particulier la fraction chlorée organique), du bois, des déchets alimentaires et du papier. Les analyses immédiates et ultimes ont été réalisées sur ces matériaux avant leur utilisation. De même, le réactif CaO a été caractérisé du point de vue de ses propriétés physico-chimiques telles que la surface spécifique, la porosité et la composition. La se tio . d́ it les ́thodes d ́ ha tillo age et d analyses. En particulier, la ́thode de aptage du gaz HCl pa a so ptio da s u e solutio asi ue, et l a al se pa chromatographie ionique sont décrites. La caractérisation des goudrons par chromatographie gazeuse couplée à la spectrométrie de masse est détaillée. Enfin, les techniques permettant de caractériser le résidu solide de pyrolyse sont expliquées : l a al se thermogravimétrique, la fluorescence X, la microscopie électronique à balayage et la i oa al se ED", la diff a tio de a o s X et l adsorption de gaz. Ces techniques pe ette t d a ́de à la o phologie, la te tu e, ais aussi à la o positio glo ale et locale.Pou o pĺte e t a ail, u e ́tude the od a i ue th́o i ue a ́t́ e ́e à l aide du logiciel FactSage 6.3. Le fonctionnement du logi iel do ́es d e t ́e et de so tie et les bases de données associées sont détaillés à la section 2.4.

	Ce chapitre présente la caractérisation des ressources et des réactifs utilisés dans cette
	étude ; les différents montages expérimentaux, les conditions opératoires, les méthodes
	d ́ ha tillo age et d a al ses ; et pou fi i la ́thodologie appli úe pou l ́tude
	d a al se de	le de ie.
	́a teu est o pa ́e, e p ́se e et e l a se e de CaO. U	o tage e ṕ imental
	couplé Py-GC/MS (Pyrolyse -chromatographie gazeuse / spectroscopie de masse) a été
	utiliś pou l ́tude sṕ ifi ue d ide tifi atio et de ua tifi atio des goud o s issus de la
	pyrolyse de composés modèles (PVC, déchets alimentaires, bois) en présence ou en
	l a se e de ́a tif CaO. E fi , l ́ uipe e t d a al se the og a i ́t i ue est d́ it,
	celui-i se a t l ́tude de o p ́he sio de la i ́ti ue de fo atio de HCl à pa ti de
	PVC.

La section 2.2 décrit les montages expérimentaux utilisés, ainsi que les conditions opératoires appliquées à chaque partie de l ́tude. Un réacteur tubulaire vertical, chauffé ́le t i ue e t, est utiliś pou la ise e oeu e des e ṕ ie es de itigatio de HCl. Da s ce cas, du gaz HCl est introduit directement dans le réacteur, et la capacité de rétention de HCl par le lit de réactif CaO est évaluée, selon différentes conditions opératoires. Dans un se o d te ps, l a atte e t de HCl est ́tudí e p ́se e de CaO et d u e sou e ǵ ́ a t le HCl in-situ. Dans ce cas, un réacteur tubulaire horizontal est utilisé, contenant du PVC ou du NaCl (sources organiques et inorganiques de chlore), et le lit fixe de CaO. En fonction des conditions opératoires de pyro-gazéification, la quantité de HCl mesurée en sortie de E fi la de iè e se tio . s atta he à d́ i e la ́thodologie et les pa a èt es ́ alús lo s de l a al se de le de ie. Il est e pa ti ulie d́tailĺ les at́go ies d i pa ts sélectionnées et les références de normalisation.

Résumé du Chapitre 3 Etude expérimentale et modélisation cinétique de la mitigation de HCl en présence du réactif CaO Le

  ut de e hapit e est d ́tudie la capacité du réactif CaO à immobiliser le gaz HCl à l i t́ ieu du ́a teu et d ́tudie le ́ a is e de ́a tio aff́ e t à la ́a tio gaz/solide. Pour se faire, les courbes de percée ont été obtenues sous différentes conditions opératoires, puis les modélisations thermodynamique et cinétique ont permis de mieux évaluer les o ditio s oṕ atoi es fa o a les à l i o ilisatio de HCl, et le ́ a is e asso í. Les principales conclusions de ce chapitre sont résumées comme suit. Dans la section 3.2, les calculs the od a i ues, à l aide du logiciel FactSage, ont été menés. Il apparaît que dans une gamme de température de 500 à 1200 K (227 à 927°C), l e thalpie li e de la ́a tio e t e CaO et HCl est ́gati e, la ́a tio est do fa o iśe dans le sens de la formation de CaCl 2 et H 2 O. Cepe da t, l aug e tatio de te ṕ atu e a

	est affecté.
	La section 3.3 décrit les résultats expérimentaux (courbes de percée) obtenus à
	différentes températures (550, 650, 750 et 850°C) en présence de CaO. En augmentant la
	température, les temps caractéristiques de percée et de saturation diminuent
	significativement : de 197,8 à 21,8 min (temps de percée à 550 et 850°C), et de 268,9 to 69,6
	min (saturation à 550 et 850°C). De plus, la capacité de rétention de HCl par CaO diminue de
	778,9 à 173,9 mg/g-CaO lorsque la température augmente de 550 à 850°C.
	Les ́sultats e ṕ i e tau o t ́t́ o pa ́s à	odèles d́ i a t l ́tape
	cinétiquement limitante de la réaction gaz/solide dans le modèle du noyau rétrécissant :
	diffusion externe de HCl, diffusion interne de HCl à travers la couche de produit en cours de
	fo atio , la ́a tio de su fa e e t e HCl et CaO. Les ́sultats o t e t ue l ́tape
	cinétique déterminante est la réaction chimique, puis la diffusion à travers la couche de
	p oduit fo ́e. E t e es	́gi es, u e zo e de ́gi e i te appa aît do t l ́te due
	dépend de la température. En effet, à basse température (550°C), le régime mixte apparaît

u effet ́faste su l ́ uili e de la ́a tio , p o o ua t u e aug e tatio sig ifi ati e de l e thalpie li e. La aleu de la o sta te d ́ uili e de la ́a tio passe ai si de p esque 7x10

16

à 500 K à environ 2x10

4

à K. Ai si la ́a tio este spo ta ́e, ais l ́ uili e pou u tau de o e sio de CaO d e i o %, ta dis u à °C, le ́gi e i te apparaît dès 2% de taux de conversion. Une proposition de mécanisme est suggérée à la figure 3.0.1.

Résumé du Chapitre 4 I flue e du a tif CaO su l' issio de HCl p ove a t de la conversion thermochimique de sources organiques et inorganiques

  et (3) la décomposition des polyènes et la cyclisation. Les pa a èt es i ́ti ues a a t́ isti ues o t ́t́ e t aits e p ́se e et e l a se e du ́a tif CaO. E l a se e de CaO, les al uls o t o t ́ ue l ́ e gie d a ti atio appa e te aut 136,5 kJ/mol, 193,7 kJ/mol, et 281,6 kJ/mol pour, respectivement, les composantes 1, 2 et 3.

	Résumé du Chapitre 5
	I flue e
	L ajout de CaO odifie de a iè e sig ifi ati e l ́ e gie d a ti atio appa ente globale, en la
	Contrairement au chapitre 3, dans lequel le gaz HCl était directement introduit dans la diminuant de 197,3 à 148,9 kJ/mol, et l ́ e gie d a ti atio de la o posa te , e
	chambre du réacteur, ce chapitre se concentre sur la génération in-situ de HCl à partir de l aug e ta t de 136,5 kJ/mol à 152,6 kJ/mol. D u poi t de ue i ́ti ue, la p ́se e de
	sources organiques (PVC) et inorganiques (NaCl). Cette ́tude ise à ua tifie l effet du lit CaO est ai si fa o a le à l i o ilisatio du hlo e da s le ́acteur, et à la pyrolyse de la
	fi e de CaO su l i matière organique	o ilisatio du HCl p oduit lo ale e t lo s de la o e sio
	thermochimique des 2 ressources sélectionnées.
	mitigation du HCl généré in-situ est meilleure à 550°C comparativement aux autres
	te ṕ atu es	,	et	°C . L effi a it́ à l a atte e t de HCl aug e te de a iè e
	significative lorsque le rapport molaire Ca/Cl est augmenté, et la taille des particules de CaO
	diminué. Pour une pyrolyse de PVC réalisée à 750°C, le taux de chlore piégé par CaO est de
	85,6% lorsque le ratio molaire Ca/Cl est fixé à 2:1.
	Enfin, la section 4.4 présente les calculs de modélisation cinétique de la conversion du
	PVC. La od́lisatio a ́t́ ́aliśe à pa ti d essais d a al se the og a i ́t i ue à

La section 4.2 présente les calculs thermodynamiques et les résultats expérimentaux e ́s e l a se e de CaO. Co e a t NaCl, seule e t , % du o te u hlo ́ est o e ti e phase gazeuse lo s de la p ol se de la essou e jus u à °C . Il a aussi ́t́ montré que le taux de conversion du chlore vers la phase gazeuse est légèrement augmenté e p ́se e de dio gè e, d eau ou de i ́ au tels ue "iO 2 . Expérimentalement, le taux de conversion maximal atteint était de 8% en combinant les 3 promoteurs à 950°C. Par o t e, la p ol se du PVC o duit à la li ́ atio de HCl dès °C. L analyse thermogravimétrique, couplée à une analyse des gaz émis par FTIR, a montré que la gamme de te ṕ atu e d ́ issio de HCl à pa ti du PVC se situe e t e et °C. Ai si, toute la fraction chlorée du PVC conduit à la formation de HCl, dans cette gamme de température, par des réactions de déchloration. Les pertes de masses observées au-delà de 400°C concernent la pyrolyse de la matière organique restante. La se tio . est d́díe à l ́tude de l effi a it́ de la ́te tio d HCl pa le ́a tif CaO lo s de la p ol se du PVC. De a iè e ǵ ́ ale, il est d́ o t ́ ue l effi a it́ à la différentes vitesses de montée en température, et en considérant une séparation du phénomène global en 3 composantes : (1) la déchloration, (2) la formation de composés aromatiques benzéniques

du a tif CaO su l' issio de goud o s lo s de la pyrolyse des composés modèles représentatifs des déchets municipaux

  Ce travail vise à étudier la capacité de CaO à limiter les goudrons émis par les composés modèles des déchets municipaux. Ainsi, les expériences sont menées avec chaque composé odèle ois, d́ het ali e tai e, PVC p is i di iduelle e t e p ́se e et e l a se e de CaO. Les goudrons émis sont identifiés et quantifiés. Ce travail a été réalisé avec le montage Py-GC/MS à différentes températures. Les conclusions principales sont résumées ci-dessous :

 La présence de CaO, lors de la pyrolyse du bois et des déchets alimentaires, favorise la formation de composés aromatiques et limite la quantité relative de composés oxygénés dans les goudrons  Le PVC p oduit p i ipale e t des d́ i ́s a o ati ues e aiso de l ́li i atio du chlore (voir figure 5.0.1). Le rôle du réactif CaO est surtout mesuré en termes d i o ilisatio de HCl, o duisa t à u e ́du tio des d́ i ́s o ga i ues chlorés Figure 5.

0.1 Mécanisme de formation des composés chlorés lors de la pyrolyse du PVC

	 A l i e se, le tau de BTXN est sig ifi ati e e t ́duit lo s de la p ol se du PVC
	en présence de CaO
	 E p ́se e de CaO, l aug e tatio de la température conduit à une décroissance
	monotone du taux de BTX entre 550 et 850°C, et à une augmentation simultanée
	du taux de naphtalène.

 La présence du réactif CaO augmente la production de goudrons à haute valeur ajoutée, tels que les BTXN (benzène, toluène, xylène et naphtalène), issus de la pyrolyse du bois et des déchets alimentaires.

Résumé du Chapitre 6 I flue e du a tif CaO su l' ission de HCl et de goudrons lors de la pyrolyse d'u la ge si ul de déchets municipaux

  Dans ce chapitre, les déchets municipaux sont simulés par un mélange modèle contenant des fractions définies de chaque composé modèle. Le ut est d ́ alue l effet du réactif CaO placé in-situ dans le réacteur et de la température sur la rétention du chlore, et sur la qualité/quantité des goudrons émis. Les résultats expérimentaux sont présentés en E fi , la se tio . o t e ue l utilisatio de CaO, ai si ue l aug e tatio de la température, permettent de réduire efficacement la quantité de goudrons produits au cours de la pyrolyse du mélange simulé. La composition des goudrons est aussi affectée par les conditions opératoires. En effet, à basse température les composés majoritaires sont de nature aliphatique, tandis que les hautes températures conduisent à la formation de composés aromatiques, par des réactions de craquage et de cyclisation. En présence de CaO, l i pa t ajo itai e e t o se ́ est la ́du tio des goud o s de atu e o ǵ ́e, et l aug e tatio de aphtalè e et de ses d́ i ́s.

	Résumé
	développant successivement : (1) le bilan matière et la composition du gaz ;	l effi a it́ à
	l a atte e t de HCl et	le a uage des goud o s.
	produit, les quantités de H 2 produites représentent 6,76 à 79,52 mL/g-mélange à
	respectivement 550 et 850°C, puis sont augmentées de 25 et 88% en présence de CaO aux
	mêmes températures.
	Les expériences menées en utilisant le réactif CaO à la section 6.2 montrent que le taux
	de résidu solide a tendance à augmenter tandis que la production de goudrons diminue.
	L aug e tatio de asse du solide provient de la rétention efficace des gaz acides tels que
	HCl et CO 2 par CaO. En effet, à la section 6.3, la rétention du chlore est évaluée à 92,4% à
	°C et , % à	°C, pou u atio olai e Ca/Cl de : . L aug e tatio du atio olai e
	de Ca/Cl améliore significativement le taux de rétention du chlore (par exemple, à 750°C, la
	rétention du chlore augmente de 42% à 86% pour des ratios molaires de 1:2 et 4:1). Les
	analyses physico-chimiques du solide après les tests montrent que le chlore est immobilisé
	dans la fraction solide, principalement sous forme de CaClOH. Ces travaux menés avec le
	mélange simulé montrent que les conditions optimales pour immobiliser le chlore et fournir
	un gaz de bonne qualité énergétique en quantités élevées sont 750°C et un ratio molaire
	Ca/Cl de 2:1.

Les expériences menées à partir de la composition simulée sont présentées à la section . . Les ́sultats o t e t ue l aug e tatio de te ṕ ature augmente le taux de gaz p oduit et do di i ue les f a tio s du ́sidu solide et des goud o s. Ai si, e l a se e de CaO, la quantité de gaz produite représente 81,9 mL/g-mélange à 550°C et 451,4 mL/g-mélange à 850°C. En présence de CaO, et aux mêmes températures, ces valeurs sont portées à 110,9 et à 510,2 mL/g-mélange. Du point de vue de la qualité énergétique du gaz

du Chapitre 7 Analyse de cycle de vie appliquée aux technologies de conversion thermochimique des déchets municipaux et évaluation des technologies de contrôle du HCl émis L

  o je tif de e hapit e est l ́ aluatio et l opti isatio des te h ologies a tuelles de traitement thermique des déchets municipaux par analyse du cycle de vie (ACV). De plus, une attention particulière a été portée aux systèmes de pyro-gazéification couplés à diverses st at́gies de d a atte e t des ́ issio s de HCl pa e e ple e -situ à basse température, ex-situ à haute température, et in-situ à la température de pyro-gaźifi atio . L o je tif de ette ́tude est d appo te des ́ĺ e ts et outils d́ isio els pou l a ́lio atio lo ale des stratégies de management des déchets municipaux. Dans ce but, deu ́tudes d ACV o t ́t́ menées.

La première étude (section 7.2) est p i ipale e t d́díe à l ́ aluatio de diff́ e ts

i -situ dans le réacteur de pyro-gazéification : impact sur l'abattement des émissions de HCl et sur la valorisation des goudrons L

  Les analyses physico-chimiques menées sur les résidus solides après abattement de HCl montrent que la surface spécifique et le volume poreux du solide diminuent sig ifi ati e e t a e l aug e tatio de la te ṕ atu e. Les i ages de i os opie ́le t o i ue à ala age o fi e t ue la te tu e po euse se de sifie d auta t plus rapidement (sous-entendu à plus faible taux de conversion de CaO) que la température est élevée. ́tude porte sur la capacité de CaO à retenir le HCl généré in-situ dans le réacteur par pyro-gazéification de ressources contenant du chlore organique ou inorganique. Typiquement pour les températures de pyrolyse envisagées (550-°C , l ́ issio de HCl provie t uasi e lusi e e t de essou es o ga i ues. L effi a it́ de la ́te tio d HCl pa le réactif CaO lors de la pyrolyse du PVC atteint 85,6% à 750°C et pour un ratio molaire Ca/Cl de 2:1. La cinétique de la pyrolyse du PVC a été étudiée au travers de 3 composantes : (1) l ́ issio de hlo e, (2) la formation de composés aromatiques benzéniques et (3) la d́ o positio des pol è es et la lisatio . L ajout de CaO odifie de a iè e sig ifi ati e l ́ e gie d a ti atio appa e te glo ale, e la di i ua t de , à , kJ/ ol, et l ́ e gie d a ti atio de la o posa te , e l aug e ta t de , kJ/ ol à , kJ/ ol. L ́ issio de HCl et so i o ilisatio pa CaO o t ́t́ ́tudíes e utilisa t u mélange simulé de résidus solides. La quantité de gaz produite augmente de 336.2 mL/g-mélange à 375.7 mL/g-mélange en présence de CaO à 750°C. La qualité énergétique du gaz produit, est aussi aug e t́e pa l ajout de CaO. A 750°C, les quantités de H 2 produites augmentent de 49.3 mL/g-mélange à 64.7 mL/g-mélange. L a ́lio atio de la ualit́ du gaz p oduit est à la fois le ́sultat de l a atte e t de polluants tels que HCl mais aussi CO 2 , et le craquage des goudrons ; augmentant ainsi le taux de conversion vers la phase gazeuse. De plus, la présence de CaO améliore la qualité de l huile de p ol se e ́duisa t la pa t de o pośs o ǵ ́s. Le meilleur compromis quantité de gaz / qualité énergétique du gaz / immobilisation de HCl / craquage des goudrons est obtenu à 750°C et pour un ratio molaire Ca/Cl de 2:1.

	 Utilisatio d'u	a tif CaO

 Analyse

de cycle de vie appliquée aux technologies de conversion thermochimique des déchets municipaux couplées à différentes options d'a atte e t des issio s chlorées L

  ́ aluatio de diff́ e ts s stè es e ista ts de t aite e t the i ue des déchets municipaux : incinération conventionnelle (S1), pyrolyse (S2), gazéification (S3), et le système de gazéification / fusion des cendres (S4) a été réalisée. Les calculs montrent que le scénario 3 (gazéification) conduit aux meilleures performances environnementales pour les catégories suivantes : le ́ hauffe e t li ati ue GWP , l a idifi atio AP et l ́ oto i it́ te est e (TETP) ; mais conduit à un impact négatif su l ́ oto i it́ a uati ue FAETP . Le s ́ a io a l i pa t le plus ́gatif su les facteurs GWP, TETP, et la toxicité humaine (HTP) ; mais en o t epa tie appa aît effi a e pou ́dui e l i pa t e i o e e tal asso í à l eut ophisatio EP et à l ́ oto i it́ a uati ue. Baś su les pe fo a es environnementales, les scénarios peuvent être classés comme suit (du meilleur au moins

Appendix

Appendix

Gasification system (S3) leads the least environmental impacts for most of the considered categories such as GWP, AP, and TETP; but it also causes the highest environmental burdens of FAETP. Besides, the environmental impact category of HTP from S3 is also relatively high and ranked 3 rd among the evaluated systems. Gasification-melting Chapter 7: Life cycle assessment of MSW thermal technologies with different dehydrochlorination options 189 system (S4) performs worst in the environmental categories of GWP, TETP, and HTP; and it also has a relatively poor behavior on environmental impact of AP, ranking 3 rd among the four systems. However, S4 system is effective to reduce the environmental impact of EP and FAETP; least environmental burdens of EP and FAETP are generated from S4 compared to the other three systems. For MSW direct incineration system (S1), the environmental impacts of AP and EP from S1 are the highest of all the systems; unfortunately, no impact category from S1 is ranked the best among the systems. Most of the environmental impacts derived from pyrolysis system (S2) are ranked in the middle level of the four systems, except for HTP; HTP from S2 is the lowest among the systems. Table 7.5 summarizes the ranking results of the impact categories. Accordingly, the environmental sustainability of each system in descending order is: S3, Gasification > S2, Pyrolysis > S4, Gasification-melting > S1,

I i e atio > ea s the fo er system performs better than the latter one). a the number represent the ranking result of the impact categories: 1 refers to the best performance and 4 represent the worst performance; b the sum of the ranking results is a reference to compare the overall performance of the systems; c the best and worst performances of the considered system on the specific impact categories are represented by bold type and marked as red and dark blue, respectively.

 Thorough and detailed life cycle assessment of pyro-gasification-based WtE systems

In this work, the comparisons between MSW pyro-gasification and incineration, as well as the potential products downstream applications, have been evaluated using LCA methodology. However, the evaluation is only based on four selected plants. Therefore, thorough and detailed comparison of all kinds of MSW thermal conversion systems should be conducted to further optimize the MSW thermal conversion systems and to guide the future application of pyro-gasification based WtE systems.

À partir des conclusions établies par ce travail, de futures recherches pourraient pe ett e d app ofo di les poi ts sui a ts :

 Amélioration des propriétés du réactif CaO pour la mitigation de HCl à haute température :

(i) Modifier les propriétés physico-chimiques pour augmenter la capacité de rétention.

(ii) Etudier la régénération ou la réutilisation du CaO usagé. Appendix II

List of Tables

Publications in international conferences

 Yuanjun TANG, Yong CHI. Exergy analysis and exergetic life cycle assessment of MSW pyrolysis and gasification. HCl dans cette couche poreuse en croissance. Pour simuler la génération de HCl in-situ par une source organique, la pyro-gazéification du PVC a été effectuée en présence ou en absence de CaO. Une modélisation des différentes ́tapes de la d́ o positio du PVC e p ́se e de CaO a o t ́ ue l ́ e gie d a ti atio apparente de la réaction de déchlorination était légèrement augmentée (de 136,5 à152,6 kJ/ ol ta dis ue l ́ e gie appa e te glo ale de la p o-gazéification du PVC est diminuée de 197,3 à 148,9 kJ/mol en utilisant CaO.

La génération in-situ de HCl a aussi été réalisée en utilisant un mélange modèle de d́ hets u i ipau , o te a t du hlo e o ga i ue et i o ga i ue. L a atte e t e HCl a été évalué, et la nature chimique des goudrons a été analysée. En présence de CaO, la part de composés organiques oxygénés est réduite, améliorant ainsi la qualité des goudrons en ue d u affi age e io-huile.

Une approche par analyse de cycle de vie a aussi été réalisée pour la pyro-gazéification et l i i ́ atio des d́ hets u i ipau . La od́lisatio a ́t́ ́aliśe e e isageant 3 scénarios de déchlorination : (1) gazéification avec fusion des cendres puis une étape de traitement du chlore ex-situ à basse température, (2) gaźifi atio sui ie d u e ́tape d ́pu atio du hlo e e -situ à haute température, (3) gazéification couplée à une déchlorination in-situ en présence du réactif CaO. Ce dernier scénario hypothétique offre les meilleures performances environnementales (par exemple: acidification, réchauffement, ́ issio s pollua tes to i ues…. ais este u odèle th́o i ue et simplifié.

Mots

Chlorinated contaminants mitigation during pyro-gasification of wastes using CaO reactant: experimental and life cycle assessment

Thermal treatment of municipal solid waste (MSW) attracts increasing attention due to the associated environmental and energy benefits. However, due to the chlorinated components of the MSW (salts, plastics) hydrogen chloride (HCl) is usually generated and may cause corrosion, toxic organic contaminants formation, acidification, etc. The present study focuses on the reactivity of a calcium oxide (CaO) reactant for in-situ mitigation of released HCl from thermal treatment of MSW.

In this work, sorption of HCl gas in CaO reactive bed has been experimentally and theoretically studied. First, it is shown that the use of CaO is effective to remove released HCl gas, by forming a CaCl 2 layer at the surface of CaO particles. However, temperature of the reactor is a key process parameter, since the removal capacity of HCl decreases significantly from 778.9 to 173.9 mg/g-CaO with the increase of temperature from 550 o C to 850 o C. A kinetics analysis has been developed by comparing experimental data with models describing the reaction at the particle surface. It has been concluded that the sorption of HCl at the CaO particle surface is firstly limited by the heterogeneous gas-solid reaction, followed by the HCl diffusion through this porous growing layer.

To simulate the in-situ generation of HCl from organic source, pyro-gasification of PVC has been performed with or without CaO addition. The experimental data have been then used for the modeling of the different PVC decomposition steps. Although the average apparent activation energy of pseudo dehydrochlorination reaction is increased from 136.5 kJ/mol to 152.6 kJ/mol with the addition of CaO, the apparent activation energy of the overall PVC decomposition has been decreased from 197.3 to 148.9 kJ/mol by using CaO reactant.

In-situ generation of HCl from organic and inorganic sources has also been conducted using simulated MSW. HCl mitigation has been evaluated together with the chemical speciation of the produced tars. Using CaO, the amount of oxygenated organic compounds has been reduced, improving the quality of the tars for a further bio-oil upgrading.