High-Temperature Gasification of Chlorinated Hydrocarbons: Thermodynamic Calculation

1. Introduction

Chlorinated hydrocarbons form the basis of a wide range of industrial wastes. They are contained in polymers and by-products of their synthesis, solvents, finishing materials, resins, dyes, and flame retardants, as well as enter the environment with wastewater, municipal and industrial waters, and landfill leachates. The neutralization of such substrates represents a critical environmental challenge for modern society. This is due to the high risks of emission of extremely hazardous chemical substances—chlorine, phosgene, furans, dioxins, and polychlorinated biphenyls—into the atmosphere, soil, and water bodies [1].

In accordance with requirements for the production and disposal technologies of organochlorine compounds, key tasks include [2] minimizing the volume of waste and by-fraction generation, involving by-products in useful practical turnover, and ensuring the complete environmental safety of technological cycles. Consequently, target recycling algorithms must guarantee the exhaustive conversion of CHCs into non-toxic components, such as hydrochloric acid, water, syngas, and carbon dioxide. At the same time, the secondary formation of hazardous ecotoxicants – dioxins, chlorine, phosgene, furans, polycyclic aromatic hydrocarbons, and soot – must be completely eliminated.

To date, an extensive range of methods has been proposed for the destruction of organochlorine waste. These include disposal in specialized repositories [3], adsorption [4], chlorolysis [5], liquid-phase oxidation [6], alkaline hydrogenation [7], dehydrochlorination [8], polycondensation [9], direct incineration [10,11], pyrolysis [12,13], plasma-chemical destruction [14,15,16], autothermal gasification with oxygen blasting [17], as well as allothermal catalytic steam conversion [18,19,20].

The most recent research emphasizes thermochemical conversion and advanced catalytic methods to address the challenges of processing chlorine-rich waste, focusing on in-situ contaminant mitigation, tar abatement, and the distribution of chlorine species during the degradation of complex feedstocks [21,22,23]. Key advancements include sorbent-based hot gas cleaning [24], the utilization of supercritical water in hydrothermal processes to suppress dioxin formation [25], and plasma-driven electrified gasification to effectively break stable C-Cl bonds [26].

Despite the obvious technological advantages of these approaches [27,28], most of them do not ensure absolute safety and completeness of the conversion of CHCs into harmless substances. The only exception is high-temperature plasma destruction [29]; however, its widespread implementation is limited by high energy intensity and significant capital costs.

A particular hazard when using thermal disposal methods is the secondary synthesis of dioxins. It is known [30] that these most dangerous by-products are actively formed at relatively low temperatures (520–670 K) in the presence of oxygen, active chlorine, catalysts, and carbon matrices (including soot). The thermal decomposition of dioxins is initiated only at temperatures above 900 K. In this regard, effective technologies for the thermal disposal of CHCs must be implemented in a temperature regime significantly exceeding 900 K, with the strict exclusion of the simultaneous presence of oxygen, active forms of chlorine, and particulate carbon in the disposal reactors.

As an affordable alternative to costly plasma technologies, the method of allothermal, high-temperature, non-catalytic steam – carbon dioxide gasification of organic waste was proposed in [31]. This technology exhibits superior energy efficiency over plasma gasification by requiring minimal electrical power, which is strictly limited to control systems, ignition, and cooling components [32]. Thermodynamic modeling of this atmospheric-pressure process for liquid hydrocarbon waste demonstrated that conversion occurs within a stream of an ultra-superheated (above 2000 K) gasifying agent (GA) – a mixture of steam and carbon dioxide [33]. The high-temperature GA is sourced from the atmospheric-pressure expansion products of pulsed or continuous detonation of an oxygen/air mixture with methane or self-generated syngas. Thermodynamic analysis of n-hexadecane processing (simulating waste motor oil) confirmed complete conversion into pure syngas composed exclusively of hydrogen (56.5 vol.% dry) and carbon monoxide (43.5 vol.% dry), yielding a lower heating value (LHV) of approximately 19.4 MJ/kg [33].

This work is a direct continuation of the research initiated in [33]. Its key objective is to determine the optimal modes of allothermal high-temperature non-catalytic steam – carbon dioxide gasification of CHCs using thermodynamic modeling. Within the scope of this task, the process parameters ensuring the complete conversion of CHCs into non-toxic useful products are established, which allows for eliminating the risks of hazardous ecotoxicant formation.

2. Materials and Methods

2.1. Chemical Compounds

The substances listed in Table 1 are selected as the objects of study in ascending order of the chlorine mass fraction ($W$) in their composition. In addition, Table 1 provides references to literature sources containing experimental data on the composition of gaseous products from the catalytic steam conversion of the respective compounds. These profile datasets were used to validate the calculation methodology.

2.2. Calculation Method

Thermodynamic modeling of the gasification process was performed using the Aspen Plus software package (version V10, 2017) [42] in a zero-dimensional approximation. The calculation algorithm is based on the Gibbs free energy minimization procedure. The following assumptions were made when constructing the model:

• The process under consideration is steady-state;
• Gas components are described by the ideal gas equation of state;
• Mixing of the feedstock and the gasifying agent occurs instantaneously and homogeneously;
• Chemical reactions occur instantaneously, reaching thermodynamic equilibrium;
• Hydraulic pressure drops in the system are neglected;
• The mineral fraction (ash) is absent in the feedstock composition;
• Cooling of the generated syngas to standard conditions occurs instantaneously, ensuring the “freezing” of its chemical composition.

Figure 1 presents the calculation flowsheets of the gasification process for the studied chemical compounds developed in the Aspen Plus environment.

The calculation flowsheet in Figure 1a is designed to validate the mathematical model against literature data on the steam conversion of feedstock at a pressure of 0.1 MPa. It includes inlet streams of the processed feedstock, steam, and a diluent gas (nitrogen), the ratios between which are set in strict accordance with the conditions of the cited experimental works. In a continuous isothermal gasifier reactor, homogenization of the inlet streams is performed, followed by the calculation of the corresponding equilibrium chemical composition at a fixed temperature. At the outlet of the reaction volume, a stream of wet gasification products is formed and directed to the separation block. This block provides optional separation of accompanying components: water, hydrogen chloride, nitrogen, and particulate carbon (soot). The resulting dry or purified stream of gaseous products is used to validate the calculation methodology by comparing its composition with experimental data.

The calculation flowsheet in Figure 1b is developed to model the regimes of high-temperature steam – carbon dioxide gasification of CHCs, where the detonation products of a triple methane–oxygen–water vapor mixture expanded to atmospheric pressure act as the GA. The flowsheet includes the following sequential blocks and physicochemical stages:

• Inlet streams of the combustible mixture components— O₂, CH₄, and H₂O—with mass consumption of $m_{\mathrm{o}\mathrm{x}}$, $m_{\mathrm{f}}$, and $m_{\mathrm{s}}$ (kg), respectively.
• A continuous reactor modeling the generation of the high-temperature GA (predominantly H₂O and СO₂). In this block, homogenization of the initial streams is performed at atmospheric pressure, and the thermodynamically equilibrium composition of the detonation products is calculated at a pressure of 0.1 MPa and a fixed temperature in the range of 2450 to 2850 K. Since the Aspen Plus software package lacks built-in tools for the direct calculation of the Chapman–Jouguet (CJ) parameters, the boundaries of this temperature interval were set exogenously based on preliminary modeling in specialized packages, namely SDToolbox [43] and Cantera [44]. The described algorithm allows for the adequate simulation of GA generation in the form of detonation products expanded to a pressure of 0.1 MPa.
• A stream splitting block (splitter), which receives the high-temperature GA from the continuous reactor. The splitter doses exactly 1 kg of the GA for feeding into the gasifier reactor, which significantly simplifies the subsequent interpretation and comparison of the calculation results. The excess amount of the generated GA is excluded from the further calculation cycle.
• An inlet feedstock stream with a mass of $m$ (kg), which determines the specific consumption of the processed feedstock per 1 kg of the GA. The feedstock is supplied to the system at a standard temperature of $T=$ 298 K.
• An adiabatic gasifier reactor, where 1 kg of the high-temperature GA and $m$ kg of the feedstock are first mixed at atmospheric pressure, followed by the calculation of the equilibrium composition of the reaction products at 0.1 MPa and constant entropy. The total mass of the substances formed at the outlet of the adiabatic volume is 1 + $m$ (kg).
• A separator unit that splits the wet gasification products into a gaseous phase (H₂, CO, CO₂, CH₄, etc.) and condensable species. It features optional water and HCl removal to ensure the correct evaluation of the dry gas composition. Additionally, the potential for complete dissolution of the generated hydrogen chloride within the residual steam condensate is analyzed for each considered CHC.

3. Results and Discussion

3.1. Model Validation: n-Hexadecane

As the first test object for validating the developed model, n-hexadecane (n-C₁₆H₃₄) was chosen, which is commonly used as a physicochemical surrogate for waste motor oil [45]. The calculations were carried out according to the flowsheet presented in Figure 1a. The specific LHV of the initial n-hexadecane is $Q=$44.5 MJ/kg.

In addition to the main components (H₂, CO, CH₄, O₂, CO₂, and H₂O), the list of potential steam conversion products included hydrogen peroxide (H₂O₂), as well as light hydrocarbons, alcohols, and aldehydes of the С₁–С₅ series. Furthermore, unlike the baseline model [33], solid carbon in the graphite modification (C) was introduced into the phase composition of the products to simulate the formation of particulate soot.

Table 2 provides a comparison of the calculated and experimental compositions of the syngas generated during the catalytic steam conversion of n-hexadecane at atmospheric pressure. The range of investigated temperatures ($T$) was from 923 to 1246 K, and the steam-to-carbon ($S$/C) ratio varied from 2.7 to 4.0. The table also indicates the corresponding literature sources that served as the basis for model validation. The comparison of calculated and measured concentrations was performed exclusively for long-term experiments in which the H₂, CO, CO₂, and CH₄ profiles remained stable throughout the entire reactor operation time.

The high convergence of theoretical and experimental data indicates that, under the conditions considered, the steam conversion process reaches a thermodynamically equilibrium degree of conversion, and the developed model adequately describes the composition of the gasification products. It should be noted that when using certain catalytic systems, the complete conversion of the feedstock was not achieved in the experiments, which is explained either by the short residence time of the reactants in the reaction zone or by catalyst deactivation. The presence or absence of soot formation in the catalytic bed, recorded by the authors of the cited works, is reflected in the “Soot” column (a dash indicates the absence of relevant information in the primary source).

3.2. Model Validation: Chlorinated Hydrocarbons

As the next objects for validating the calculation methodology, the CHCs listed in Table 1 were selected. Thermodynamic modeling was carried out according to the flowsheet presented in Figure 1a. In addition to the components of the baseline model for n-hexadecane, chlorine-containing substances were included in the list of potential steam conversion products: Cl₂, HCl, Cl₂O, ClO₂, CCl₄, CHCl₃, CH₂Cl₂, COCl₂ (phosgene), as well as highly toxic polychlorinated dibenzo-1,4-dioxins (using tetrachlorodibenzodioxin, C₁₂H₄Cl₄O₂, as an example).

Table 3 provides a comparison of the calculated and experimental compositions of the syngas generated during the catalytic steam conversion of various CHCs at atmospheric pressure. The range of investigated temperatures ($T$) was from 973 to 1023 K (700–750 $°C$), and the steam-to-carbon ($S$/C) ratio varied widely from 2.0 to 20.0.

3.3. Conditions for Achieving Complete Conversion into Non-Toxic Useful Products: n-Hexadecane

To determine the conditions for achieving complete conversion of feedstocks into non-toxic industrially useful components using allothermal high-temperature non-catalytic steam – carbon dioxide gasification, the conversion of n-hexadecane was examined in the first stage. Modeling of this process was carried out according to the calculation scheme presented in Figure 1b. The methodology for processing and interpreting the obtained results, described in detail below using n-hexadecane as an example, is an end-to-end approach and is applied to all CHCs investigated in this work.

Figure 2 shows characteristic calculated dependences of the process temperature and the equilibrium composition of n-hexadecane gasification products on the specific feedstock consumption m. The modeling was performed for the conditions of GA generation during the detonation of a stoichiometric methane–oxygen mixture (fuel equivalence ratio $\mathsf{Ф}=$ 1) without the addition of water vapor ($\alpha =$ 0). The point $m=0$ on the graphs characterizes the temperature and the initial equilibrium composition of the GA at the inlet of the adiabatic gasifier reactor. According to the calculations, at the inlet to the reaction zone, the GA has a temperature of $T_{ГА}=$ 2852 K and contains 56 vol.% H₂O, 19 vol.% СO₂, 12 vol.% CO, 8 vol.% O₂, and 5 vol.% H₂ (the presence of radicals such as H, O, and OH in the expanded detonation products was not taken into account during the graphical visualization of the results in Figure 2).

As follows from the graphical dependences in Figure 2, the introduction of feedstock ($m>0$) into the gasifier reactor initiates a complex of physicochemical transformations that change the temperature and component composition of the medium. In the range of specific consumption 0 $<m\le$ 0.03, a moderate increase in the temperature of the gasification products up to a maximum extremum is observed, after which it monotonically decreases at $m>$ 0.03. The initial heating of the system is due to the total exothermic effect of the reactions occurring predominantly through the oxidation of feedstock fragments by residual free oxygen in the GA. The subsequent temperature drop is determined by the dominance of endothermic reactions of steam (С + H₂O = CO + H₂) and carbon dioxide (С + СO₂ = CO + CO) conversion of carbon contained in the initial feedstock.

Thermodynamic analysis indicates the existence of a certain range of $m$ values within which free oxygen ([O₂] = 0), particulate carbon ([С] = 0), and hydrocarbons ([СН₄] = 0) are simultaneously absent from the equilibrium product composition. For the conditions reflected in Figure 2, this interval is localized within the boundaries 0.11 $\le m\le$ 0.35. This range is defined as the operating section of the gasification process, and its left and right boundaries are defined as the lower limit (LL) and upper limit (UL) of the operating section, respectively. For a more detailed analysis, an extended upper limit (EUL) is also introduced into consideration, fixing the region where [O₂] = 0 and [С] = 0, but the presence of methane is allowed ([СН₄]$\ne$0). The efficiency of the gasification process is traditionally evaluated using a set of integral indicators. These include the specific dry gas yield (DGY), carbon conversion efficiency (CCE), cold gas efficiency (CGE), and net energy efficiency (NEE).

The specific dry gas yield (DGY, nm³/kg) characterizes the volumetric yield of gaseous products per unit mass of the processed feedstock:

$$DGY={\displaystyle \frac{q_g}{m}} ({\mathrm{nm}}^3/\mathrm{kg})$$

(1)

where $q_g$ is the volume of the obtained dry syngas, nm³; $m$ is the mass of the initial feedstock, kg.

The carbon conversion efficiency (CCE, %) represents the ratio of the total mass of carbon within the target components of the dry syngas to its total amount supplied to the gasifier reactor with the feedstock and the GA:

$$CCE=\left(Y_{\mathrm{С}0}^{-1}\sum _{i=CO,{CO}_2,{CH}_4,\dots }{Y}_{\mathrm{С}i}\right)100\mathrm{\%}$$

(2)

where $Y_{Сi}$ is the mass fraction of carbon in the $i$-th component of $Y_{С0}$ is the total mass fraction of carbon entering the reactor as part of the feedstock and the GA.

The cold gas efficiency (CGE, %) determines the fraction of the feedstock’s chemical energy transferred into the target gaseous energy carrier and describes the potential thermodynamic efficiency of the process:

$$CGE={\displaystyle \frac{q_g{Q}_g}{mQ}}100\mathrm{\%}$$

(3)

where $Q_g$ is the lLHV of the obtained dry syngas, MJ/nm³; $Q$ is the LHV of the initial feedstock, MJ/kg.

The net energy efficiency (NEE, %) characterizes the ratio of the total chemical and thermal energy of the generated stream to the total energy inputs required to implement the process, including the potential of the feedstock and external energy supplies for GA generation [46]:

$$NEE={\displaystyle \frac{q_g{Q}_g+m_g∆H_g+m_v{Q}_v}{mQ+m_{\mathrm{f}}{Q}_{\mathrm{f}}+m_{\mathrm{o}\mathrm{x}}{Q}_{\mathrm{o}\mathrm{x}}}}100\mathrm{\%}$$

(4)

where $m_g$ is the mass of the obtained dry syngas, kg; $∆H_g$ is the enthalpy change of the syngas upon its cooling to normal temperature, MJ/kg; $m_v$ is the mass of steam within the gasification products, kg; $Q_v$ is the specific latent heat of steam condensation (2.26 MJ/kg); $Q_{\mathrm{f}}$ is the LHV of methane (50 MJ/kg); $Q_{\mathrm{o}\mathrm{x}}$ is the specific energy consumption for oxygen production (1.1 MJ/kg [47]).

Unlike the CGE criterion, the NEE indicator provides a more comprehensive coverage of the system’s energy balance, since it takes into account both the thermal potential of cooling the syngas and the direct energy consumption for the generation of the high-temperature GA.

Calculations similar to the baseline case (see Figure 2) were performed for GAs generated during the detonation of fuel-lean ($\mathsf{Ф}=$ 0.75) and fuel-rich ($\mathsf{Ф}=$ 1.5) methane–oxygen mixtures. The modeling was conducted both for the initial compositions without dilution ($\alpha =$0) and with the introduction of an additional amount of water vapor into the combustible mixture ($\alpha =$ 20 and 40 vol.%). Table 4 shows five calculation modes in terms of the parameters of the initial combustible mixtures ($\mathsf{Ф}$ and $\alpha$), as well as their corresponding calculated equilibrium temperatures ($T_{\mathrm{G}\mathrm{A}}$) and GA component compositions. To assess the effect of detonation product dissociation, the calculations were performed in two variants: excluding and including the presence of free radicals (H, O, and OH). A comparison of the obtained data shows that the total content of radical components in the detonation products expanded to atmospheric pressure can reach 10 vol.%.

Table 5 systemizes the results of modeling n-hexadecane gasification processes, performed both excluding and including the presence of radical components (H, O, and OH) within the expanded detonation products. For each mode, the values of the specific feedstock consumption m at the LL, UL, and EUL of the gasification operating section (referred to as the operational points (OP)), the final adiabatic syngas temperature $T_g$, the equilibrium composition of the wet syngas, as well as the equilibrium composition, volume, and mass of the dry syngas obtained after the separation of condensed moisture are given. In addition, the table presents the calculated LHVs $Q_g$ of the dry syngas and the integral process efficiency criteria: DGY, CCE, CGE, and NEE. When calculating the enthalpy term $m_g∆H_g$ in Eq. (4), the heat capacity values of the syngas components were taken at $T=$ 298 K.

Analysis of the data in Table 5 shows that complete conversion of the feedstock carbon (CCE = 100%) is ensured in all investigated modes. The specific dry gas yield at the upper limits (UL and EUL) of the operating section is expectedly higher than at the LL, reaching a level of 7.7 nm³/kg. The cold gas efficiency exceeds the 100% barrier across almost the entire range of parameters, peaking at 174%, which indicates a significant contribution of the detonation GA to the chemical potential of the products. The net energy efficiency indicator varies predominantly within the range of 85–95%, and also overcomes the 100% threshold in certain modes. The inclusion of radical species (H, O, and OH) in the GA generally alters the dry-gas performance under identical baseline conditions, exerting the maximum impact on dry-gas volume and mass (up to 15%), LHV (up to 15%), as well as CGE and NEE (up to 8%), which stems from elevated final adiabatic syngas temperatures ($T_g$) and enhanced dry-gas H₂ and CO content.

With an increase in the parameter $m$ from the LL to the EUL, an increase in the LHV of the generated syngas up to 20 MJ/kg is recorded, while the total volume fraction of combustible components in the dried and СО₂-purified stream reaches 97%. It should be emphasized that removing the rigid restriction on the presence of methane in the final gas composition (the transition from UL to EUL) provides a pronounced improvement in all techno-economic and energy indicators of the considered technological scheme.

3.4. Conditions for Achieving Complete Conversion into Non-Toxic Useful Products: Chlorinated Hydrocarbons

To determine the conditions for achieving complete conversion of CHCs into non-toxic and industrially useful components using allothermal high-temperature non-catalytic steam – carbon dioxide gasification, the conversion of chlorohexadecane was examined in the next stage. Modeling of the gasification process was carried out according to the calculation scheme presented in Figure 1b. The methodology for processing the obtained data completely corresponds to the algorithm described above for the baseline hydrocarbon feedstock (n-hexadecane); however, the composition of potential gasification products was expanded to include chlorine-containing substances.

Figure 3 shows the calculated dependences of the process temperature and the equilibrium composition of chlorohexadecane gasification products on the specific feedstock consumption m. The conditions for the generation of the high-temperature GA are identical to the baseline case (see Figure 2): detonation of a stoichiometric methane–oxygen mixture ($\mathsf{Ф}=$ 1, $\alpha =$ 0). Accordingly, at the point $m=0$, the parameters of the material and heat flows at the inlet of the adiabatic gasifier reactor completely match the data presented in Figure 2.

The main difference between the thermodynamic profiles in Figure 3 is the appearance of hydrogen chloride (HCl) in the gasification products at $m>0$, the volume fraction of which increases monotonically as the consumption of the processed feedstock increases. It should be emphasized that within the investigated range of conditions, other chlorine-containing compounds (including molecular chlorine Cl₂, phosgene COCl₂, and highly toxic dioxins C₁₂H₄Cl₄O₂ are virtually absent from the gas phase — their equilibrium volume fractions do not exceed the threshold value of 10^-6 (1 ppm).

The behavior of temperature and concentrations of the main gas components (H₂, CO, CO₂, H₂O) is qualitatively equivalent to the dependences in Figure 2. Similarly, the operating section of the gasification process stands out on the graphs, characterized by the simultaneous fulfillment of the conditions [O₂] = [C] = [СН₄] = 0. For chlorohexadecane, the operating section is localized within the boundaries 0,13 $\le m\le$ 0,41, where the LL and UL correspond to specific feedstock consumption values of $m=$ 0,13 and $m=$0,41. The extended operating section ([O₂] $=$ [С] $=$ 0 at [СН₄]$\ne$0) is limited to the range of 0,13 $\le m\le$ 0,46, fixing the EUL at $m=$ 0,46.

Similar calculations were performed for all CHCs listed in Table 1, utilizing the detonation products of a stoichiometric, dry methane–oxygen mixture ($\mathsf{Ф}=$ 1, $\alpha =$ 0) as the GA. The thermodynamic modeling evaluated two scenarios: with and without the inclusion of radical species (H, O, and OH). The results are summarized in Table 6. For each investigated compound, the table provides the chlorine mass fraction in the molecule ($W$), the specific feedstock consumption ($m$) at the operational points (LL, UL, and EUL), and the final adiabatic gasification temperature ($T_g$). Additionally, it details the wet gas composition, the volume and mass of the chlorine-free dry gas obtained after moisture and hydrogen chloride separation, the LHV ($Q_g$), and the integral efficiency indicators (DGY, CCE, CGE, and NEE). The efficiency indicators are determined based on the chlorine-free dry gas composition, where the enthalpy term ($m_g∆H_g$) in Eq. (4) is evaluated using the heat capacities of the dry gas components at $T=$ 298 K. It should be noted that during the gasification of CHCs with a high initial chlorine content, it is advisable to use the technological approach applied in [20]. It consists in the co-feeding of organochlorine wastes with aliphatic or oxygen-containing hydrocarbons, which allows for reducing the total mass fraction of chlorine in the reactive mixture. In Table 6, this method is implemented using the example of a blended feedstock consisting of chloroform and ethyl acetate (CHCl₃ + C₄H₈O₂).

From the analysis of the data in Table 6, it follows that complete conversion of the feedstock carbon (CCE = 100%) is ensured in all investigated modes of CHC gasification. The volumetric yield of dry gas at the UL and EUL of the operating section is expectedly higher than at the LL, reaching a level of 5.7 nm³/kg. The cold gas efficiency of the process exceeds the 100% barrier across almost the entire range of parameters, reaching a maximum value of 138%. The inclusion of radical species (H, O, and OH) in the GA generally yields a trend consistent with the previously outlined chlorohexadecane gasification data.

With an increase in the specific feedstock consumption $m$ from the LL to the EUL, a monotonic increase in the LHV of the generated gas up to 17 MJ/kg is recorded. At the same time, the maximum total volume fraction of combustible components in the dried and CO₂-purified product stream reaches 99%. It should be emphasized that removing the rigid restriction on the presence of methane in the final gas composition (the transition from UL to EUL) provides a pronounced improvement in all integral techno-economic and energy indicators of the considered technological scheme for CHC utilization.

The potential for complete hydrogen chloride dissolution within the residual steam condensate is also addressed. To evaluate this, Table 7 details the thermodynamic parameters governing HCl absorption by the aqueous phase of the wet gasification products. As a limiting condition, the formation of a saturated hydrochloric acid solution with a molar ratio of $\gamma =$HCl/H₂O $=$0.36 at a temperature of $T=$ 293 K and a pressure of $p=$ 0.1 MPa is adopted [48]. The specific feedstock consumption corresponding to the formation of such a saturated solution is designated as $m_{\mathrm{s}}$. As before, the GA consisted of the detonation products of a stoichiometric, water-vapor-free methane–oxygen mixture ($\mathsf{Ф}=$ 1, $\alpha =$ 0). Two calculation scenarios were compared: including and excluding radical components (H, O, and OH).

As follows from Table 7, the entire amount of hydrogen chloride generated during the gasification of all considered chlorinated compounds can be dissolved in the condensate of the residual steam. The threshold value $m_s$ is either equal to or less than the corresponding $m$ value at the EUL for a given feedstock. For certain chlorinated compounds (C₂Cl₄, C₂HCl₃, C₂H₃Cl₃, CH₃Cl), $m_s<$LL, meaning that feedstock gasification would require the presence of free oxygen in the GA, which is unacceptable [30].

Under these circumstances, two solutions are possible. The first is the external addition of steam to the wet gasification products upstream of the separator in Figure 1b to ensure complete dissolution of the generated HCl in water. The second is the co-feeding of the organochlorine feedstock with aliphatic or oxygen-containing hydrocarbons, which reduces the total mass fraction of chlorine in the reactive mixture. The effectiveness of the latter approach is demonstrated using the CHCl₃ + C₄H₈O₂ mixture as an example (see Table 7).

5. Conclusions

In this work, using thermodynamic modeling in the Aspen Plus software package, the fundamental feasibility and high efficiency of applying allothermal high-temperature non-catalytic steam – carbon dioxide gasification for the environmentally safe utilization of CHCs have been substantiated. Detonation products of ternary methane–oxygen–water vapor mixtures expanded to atmospheric pressure were considered as the steam – carbon dioxide gasifying agent.

Based on the obtained calculated data, the following key conclusions have been formulated:

Methodology validation. Validation of the developed zero-dimensional calculation model was performed using literature data on the catalytic steam reforming of pure hydrocarbon feedstock (n-hexadecane) and various types of CHCs at atmospheric pressure. Good agreement between the theoretical and experimental equilibrium composition of the syngas was demonstrated, which confirms the adequacy of the chosen thermodynamic basis (Gibbs free energy minimization).

Environmental safety. It was found that when using a high-temperature GA (with an initial temperature of $T_{\mathrm{G}\mathrm{A}}=$ 2450–2850 K) within the investigated range of conditions, the chlorine heteroatom is bound exclusively into hydrogen chloride (HCl). The equilibrium volume fractions of other chlorine-containing ecotoxicants (including molecular chlorine Cl₂, phosgene COCl₂, and highly toxic dioxins C₁₂H₄Cl₄O₂ in the reaction zone do not exceed the threshold value of 10^-6 (1 ppm), which completely solves the problem of harmful emissions during thermal waste processing.

Determination of operating sections. For each investigated feedstock, the boundaries of the operating sections were determined in terms of the specific feedstock consumption m (per 1 kg of GA), within which free oxygen, particulate carbon (soot), and hydrocarbons are simultaneously absent from the conversion products. It was shown that allowing the presence of methane in the final products (the transition to the extended upper limit — EUL) significantly improves all integral characteristics of the process.

High energy efficiency. Complete conversion of the feedstock carbon (CCE = 100%) is ensured in all operating modes of CHC gasification. The volumetric yield of dry syngas reaches 5.7 nm³/kg of feedstock at its LHV of up to 17 MJ/kg (with the total fraction of combustible components up to 99%). The cold gas efficiency exceeds the 100% barrier (up to 138%), proving the effective transformation of detonation product energy into the chemical energy of the syngas.

Waste-free chlorine separation. A chlorine capture algorithm was substantiated, based on the dissolution of generated HCl within the residual steam condensate upon cooling the gasification products to $T=$ 293 K. For highly chlorinated feedstocks (chlorine mass fraction above 70%), a moisture deficit coupled with an oxygen excess in the GA makes standard gasification unfeasible. To address this, a co-feeding method using external hydrocarbons was proposed and validated (demonstrated via a CHCl₃ + C₄H₈O₂ mixture). This approach ensures complete HCl dissolution across the entire operating range, yielding a commercially viable industrial product: an aqueous hydrochloric acid solution.

The proposed method of detonation steam – carbon dioxide gasification of CHCs is a promising, economically affordable, and less energy-intensive alternative to high-temperature plasma-chemical technologies for toxic industrial waste utilization.