One of the primary drivers behind stringent statutory frameworks for livestock by-products is the rapidly expanding volume of accumulated waste. Mismanagement of manure and droppings through storage on unequipped land sites frequently results in environmental contamination [1,2]. Within this category, swine manure (SM) poses the greatest environmental hazard because of high concentrations of toxic components, namely ammonia, hydrogen sulfide, and mercaptans [3,4]. Therefore, viable technologies for the large-scale, high-throughput processing of SM are needed, as current solutions remain largely in the development phase.

A variety of modern methods are employed for SM processing, including biochemical routes such as accelerated composting [5,6], fertilizer production [7,8], and anaerobic digestion [9-11], as well as thermochemical technologies such as direct combustion [12-15], pyrolysis [16-19], hydrothermal conversion (liquefaction [20], carbonization [21], and supercritical water gasification [22,23]), and conventional gasification [24-27].

Thermochemical methods are currently considered the most promising. In contrast to biochemical pathways characterized by prolonged processing times (weeks) and complex process control, thermochemical methods rely on relatively rapid thermolysis and conversion mechanisms. These processes effectively destroy pathogens [28], significantly reduce feedstock volume, and yield valuable outputs, with recovered products including thermal energy, liquid biofuels, and syngas or biogas (mixtures of H₂, CO, light hydrocarbons, and CO₂) suitable for renewable power generation [29]. Additionally, the process yields a highly porous solid residue with an elevated concentration of nutrients such as phosphorus and potassium [30], which can serve both as an agricultural fertilizer and a sorbent for heavy metals or organic pollutants [31]. However, thermochemical methods for SM processing are typically hindered by high moisture and ash content, requiring additional separation and drying operations.

Direct combustion of SM in grate furnaces or fluidized bed reactors offers a practical approach for energy generation and greenhouse gas mitigation [32,33]. Nevertheless, the intensive flue gas cleaning required to remove hazardous emissions, including furans and dioxins, renders this method cost-prohibitive.

Pyrolysis of SM or its compost is highly energy-intensive, typically requiring the combustion of either the SM feedstock itself or auxiliary organic matter. This process generally yields three product phases: biogas (10–20 wt.%), char (15–25 wt.%) [34], and pyrolysis oil (60–75 wt.%) [35,36]. After water separation, the higher heating value (HHV) of pyrolysis oil from raw SM and SM compost reaches 28.5 MJ/kg [37] and 31.2 MJ/kg [38], respectively. These values are significantly higher than those of wood-sawdust pyrolysis oil (23.9 MJ/kg) [38]. To upgrade this pyrolysis oil into commercial biofuel, multi-stage catalytic hydrotreating under high pressure is required [39]. However, the resulting biofuel is highly corrosive, and converting it to syngas via steam reforming is severely limited by catalyst poisoning and coking [40]. Concurrently, char from SM pyrolysis has high phosphorus, nitrogen, and sulfur contents, along with low pH and electrical conductivity (EC) values [41].

Hydrothermal methods operate under extremely high pressures, which severely restrict their widespread industrial adoption [20-23]. For instance, hydrothermal liquefaction (depolymerization) of solid organic fractions in SM uses water at pressures of 10–25 MPa and temperatures of 550–640 K to yield a primary liquid biofuel with an HHV close to that of crude oil, along with gas and solid by-products. Hydrothermal carbonization of SM occurs in an oxygen-deficient environment using a catalyst and water/steam at pressures up to 10 MPa and temperatures of 720–840 K; this pathway produces hydrogen-rich biogas and a liquid biofuel rich in phenols, nitrogen, and aromatics. Similarly, catalytic or non-catalytic gasification of SM with supercritical water (pressures > 22.1 MPa, temperatures > 650 K)—which benefits from the enhanced solubility of organic substances—yields high-quality syngas but also an aqueous residual solution containing organic pollutants (phenols, substituted phenols, carbocyclic and N-heterocyclic compounds, benzene, and substituted benzenes).

For conventional SM gasification [24-27], a gasifying agent (GA) such as air, steam, or carbon dioxide is introduced at high temperatures (above 800 K) under atmospheric pressure. In an oxygen-deficient or completely oxygen-free environment, the organic constituents of SM are converted into syngas or off-gas. This gas can be used for chemical synthesis (hydrogen, ammonia, methanol, etc.) or combusted in boilers, gas engines, and gas turbines. Gasification is more energy-efficient than pyrolysis [42]. Although pyrolysis oil is more energy-dense than gasification products, the syngas obtained via gasification still maintains a substantial HHV of 15–20 MJ/Nm³ [24,26].

Utilization of steam and carbon dioxide as GAs offers distinct advantages. First, because steam and CO₂ comprise only H/O and C/O atoms, the resulting syngas is free from inert gas dilution. Second, the high enthalpy of H₂O and CO₂ significantly reduces the required GA volume. Third, operating in the absence of free oxygen prevents the formation of toxic compounds such as furans and dioxins [43], thereby simplifying downstream purification. Fourth, steam gasification yields several times more hydrogen than air-blown gasification. Finally, CO₂ can be sourced directly from the flue gases of thermal power plants, providing a viable pathway for carbon capture and footprint reduction.

In general, steam/CO₂ gasification involves numerous heterogeneous and homogeneous endothermic and exothermic reactions between active radicals, atoms, and molecules, as well as electronically excited and ionized species in plasma gasification. A highly simplified set of these overall reactions can be found, for instance, in [44]. Recent literature indicates that H2O conversion proceeds via the highly endothermic primary steam-char gasification pathway, which is strongly coupled with gas-phase water-gas shift kinetics and steam reforming of hydrocarbons to maximize H₂ yields. Concurrently, at high temperatures, CO₂ reaction pathways are dominated by the heterogeneous Boudouard reaction with char, alongside homogeneous pathways such as the reverse water-gas shift and dry reforming of methane and tars. In practice, complex kinetic modeling is frequently substituted with equilibrium calculations, which generally predict qualitative trends rather than exact temperatures and species concentrations. Specifically, thermodynamic analyses of biomass steam/CO₂ gasification demonstrate that at temperatures around 900 K, carbon and oxygen are bound in CO₂, tar, and char, indicating low conversion rates. Above 1200 K, CO₂ reacts with the available carbon to yield CO, while oxygen preferentially reacts with carbon to form CO and CO₂ rather than with H₂ to form water. Finally, at temperatures exceeding 1500 K, tar and char are completely converted into syngas consisting primarily of H₂ and CO. Discrepancies between equilibrium predictions and experimental data typically stem from thermal losses, incomplete mixing, finite heat and mass transfer rates, and limitations imposed by chemical kinetics.

Among thermochemical methods, gasification is considered the most attractive for SM processing. However, conventional low-temperature technologies (up to 1200 K) have several major drawbacks, including incomplete carbon conversion, high concentrations of condensable hydrocarbons (tars) in the off-gas, and the resulting need for tar and char utilization, which can hinder industrial-scale implementation [1,44]. To overcome these limitations, high-temperature gasification can be conducted at GA temperatures exceeding 1500 K. At these elevated thermal levels, the process yields clean, high-quality syngas alongside a stabilized, non-hazardous mineral slag.

The most prominent high-temperature methods include allothermal electric-arc [45] and microwave [46] plasma gasification. In electric-arc reactors, the GA passes through an electric discharge plasma, reaching temperatures of ~10,000 K, and reacts with the waste at 1500–2000 K to convert it into syngas and slag, whereas in microwave reactors, the feedstock is heated via high-frequency (several GHz) electromagnetic radiation, driving gasification reactions within distributed “microplasma” nodes at temperatures up to ~5000 K. Despite these performance advantages, plasma gasification technologies face economic and operational barriers, namely intensive electricity consumption and the requirement for specialized, expensive refractory materials [44].

As an alternative to conventional allothermal processes, the pulsed detonation gun (PDG) gasification method [47] has proven effective for treating diverse organic wastes [48-51]. The PDG-generated steam/CO₂ mixture reaches temperatures over 2000 K, which is on par with electric-arc plasma gasification, but this approach also circumvents the major drawbacks of plasma systems outlined above.

This study evaluates the application of this high-temperature technology to SM with varying moisture contents ( $\alpha =70\mathrm{\%},45$ %, and 15%) to identify the key criteria required for scaling the system into a viable and sustainable industrial process by examining syngas quality, the impact of biomass pre-drying, solid residue properties, thermal and mass balances, cold gas efficiency, as well as computational model validation, sustainability, and process scale-up.

2.1. Feedstock and Gasifying Agent

Raw wet SM served as the primary feedstock. Its solid fraction comprised irregularly shaped particles with an average size of approximately 1 mm, while embedded bedding material ranged from 10 to 50 mm. The bulk density of the raw SM samples ranged from 600 to 1250 kg/m³. Hot-surface drying characterization indicated that the moisture content of the raw wet SM spanned 67–74 wt.% (averaging 70 wt.%). The natural gas (NG) utilized as the fuel for the PDG consisted of 96.1 vol.% CH₄, 2.1 vol.% C₂H₆, 0.6 vol.% C₃H₈, 0.2 vol.% C₄H₁₀, and 1.0 vol.% N₂.

2.2. Experimental Setup

Figure 1 shows a schematic of the experimental setup, which primarily comprises a PDG, a vertical flow reactor, a gas cleaning system, and an off-gas burner. The PDG consists of a 600-mm-long cylindrical tube with an internal diameter (ID) of 50 mm, which is connected to a 900-mm-long splitter tube (40 mm ID). This assembly includes a mixing and ignition device and a cooling jacket. To determine the detonation velocity within the PDG, two ionization probes (IPs) are positioned 250 mm apart. The PDG operates at a frequency of $=$ 0.7–1 Hz, injecting a high-velocity (~1 km/s), high-temperature (~2000 K) GA directly into the flow reactor (indicated by the dashed red arrow in Figure 1).

The flow reactor consists of a 50-L vessel fabricated from a standard gas cylinder shell. Unlike the design in [51], this reactor features an upper hatch for feedstock loading and a bottom flange for tangential connection to the PDG via the splitter tube (Figure 2). The discharge line for gasification products (comprising off-gas, solid residues, and unreacted GA) exits through the upper reactor flange. These products flow continuously out of the reactor via a central 15-mm tube that extends 300 mm downward into the reactor cavity from the upper reactor flange (indicated by the solid red arrow in Figure 1).

Feedstock can be introduced into the system using two configurations: (1) batch feeding via the upper hatch (baseline operating mode), and (2) continuous feeding through the lower hatch. For continuous operation, a custom 3-L hydraulic piston feeder based on a concrete-pump principle was developed. This mechanism ensures a stable feedstock delivery rate that is virtually independent of external operating factors.

The gasification process entails entraining fine solid particles from the reactor bed, with carbon conversion degrees varying across a wide range. To manage this, the downstream product purification system comprises a gas cooling unit sandwiched between two cyclones, culminating in an air-blown diffusion burner with a pilot flame for off-gas flaring and visual process monitoring. The product stream first passes through a 5-L dry cyclone for particulate separation, then through the cooling system, and finally into a 2-L wet cyclone to capture the remaining condensate.

Mechanistically, the gasification process in the flow reactor shares key thermodynamic and transport features with combustion in a compression-ignition (Diesel) engine:

• -

  Owing to external cooling, the reactor wall temperature remains substantially lower than that of the GA stream.

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  Chemical transformations occur predominantly within the feedstock–GA mixture inside the reactor core at elevated temperatures.

• -

  Due to the endothermic nature of the gasification reactions, the exiting off-gas temperature is significantly lower than the initial GA temperature.

• -

  Each operational cycle involves the dilution of the fresh feedstock–GA charge with residual off-gas inside the reactor volume.

Appendix A details the specific diagnostic and measurement techniques used in this study.

2.3. Experimental Methodology

The feedstock was introduced into the reactor using either batch or continuous modes. In the batch configuration, a 0.5–1.0 kg charge of SM was loaded into the flow reactor prior to each test. Subsequently, the data acquisition, gas cooling, and PDG subsystems were activated, and the PDG cyclic operation was initiated at a designated frequency. The gasification process comprised two distinct stages: (i) moisture heating and evaporation, and (ii) thermochemical conversion of the dried biomass. The transition between these stages was monitored by reactor temperature and the presence of off-gas combustion at the diffusion burner; cessation of combustion marked the nominal completion of the gasification process. Although the recorded gas and wall temperatures in the upper reactor zone dropped to 250–350 °C by the end of the process, the primary gasification reactions occurred at localized, instantaneous GA temperatures exceeding ~2000 K. To investigate thermal effects, certain batch experiments involved preheating the empty reactor to a wall temperature of 170–300 °C. The batch tests evaluated SM with high (α ~ 70%), medium (α ~ 45%), and low (α ~ 15%) moisture contents.

Continuous feedstock delivery was executed via the hydraulic piston feeder at a volumetric flow rate of 0.5–3.3 mL/s, injecting into either an unheated reactor or one preheated to a wall temperature of 400 °C. Preheating increased the average internal temperature and improved the uniformity of interaction between the feedstock particles and the supersonic GA jets, stabilizing the upper wall temperature at approximately 400 °C. To minimize parasitic energy costs from moisture evaporation, the piston feeder mechanically dewatered the loose, wet feedstock before injection. This compaction stage expressed excess water, reducing the feedstock density nearly threefold from 0.7 to 0.25 g/mL.

2.4. Methodology of Thermodynamic Calculations

In the thermodynamic model, the fuel (methane) and oxidizer (oxygen) are supplied to the PDG, premixed, and charged into the detonation tube. Upon ignition, a detonation wave propagates through the mixture, converting it into detonation products composed primarily of high-pressure, high-temperature steam and carbon dioxide. As these products enter the flow reactor, they expand to 0.1 MPa, accompanied by a corresponding temperature drop. The dry SM introduced into the reactor undergoes thermal, mechanical, and chemical treatment by the dense (~2 kg/m³) jet of detonation products, converting the biomass into off-gas. This off-gas continuously exits the reactor and is routed to downstream applications after cooling.

Thermodynamic calculations were performed using the methodology detailed in [52] under adiabatic conditions, neglecting heat losses to the environment. The GA is assumed to comprise the products of an ideal detonation of a stoichiometric methane–oxygen mixture, fully expanded to 0.1 MPa. The mixing between the feedstock and the GA is assumed to be instantaneous, homogeneous, and complete, with graphite representing the condensed phase in the resulting gasification products.

3.1. Measurements of Feedstock Properties

The average elemental composition of the dry SM, characterized using a PE 2400 Series II automatic CHNS/O analyzer (PerkinElmer Instruments, Shelton, CT, USA), comprised 45.71 wt.% C, 37.72 wt.% O, 5.78 wt.% H, 1.64 wt.% N, 0.59 wt.% S, and 8.58 wt.% ash. The hydrogen content in the dry SM closely matches that of cellulose (C₆H₁₀O₅)_n, whereas the C and O concentrations deviate upward and downward, respectively, relative to cellulose (which contains 44.44% C, 49.34% O, and 6.22% H). Consistent with this, a comparison of the IR spectra of cellulose and the SM samples, recorded on an IRTracer-100 spectrometer (Shimadzu, Japan), revealed strong structural similarities. This indicates that the feedstock includes a cellulose-rich bedding material, such as wood sawdust. Beyond C, O, and H, the biomass contains minor nitrogenous and sulfurous compounds. Elemental X-ray fluorescence (XRF) analysis, performed on an ARL PERFORM’X spectrometer (Thermo Fisher Scientific Inc., USA; X-ray tube power 2500 W), demonstrated that the mass fraction of elements with an atomic mass greater than 11 is 6.8%. The inorganic matrix is primarily composed of Si, Ca, P, Mg, K, S, Na, Cl, Al, and Fe compounds. Finally, bomb calorimetry using an ABK-1V system (Russia) quantified the higher and lower heating values of the dried SM at $Q_{\mathrm{H},\mathrm{S}\mathrm{M}}$ = 16.5 MJ/kg and $Q_{\mathrm{L},\mathrm{S}\mathrm{M}}$ = 14.6 MJ/kg, respectively (see Appendix B).

3.2. Calculated Chemical Composition of Gasification Agent

The GA was generated via the pulsed detonation of a stoichiometric NG–oxygen mixture with a fuel-to-oxidizer equivalence ratio of $\mathrm{Ф}\approx$ 1.0. Given that methane is the predominant component of NG, the fuel was modeled as pure methane in the thermodynamic simulations. The computed Chapman–Jouguet (CJ) detonation velocity for the stoichiometric methane–oxygen mixture reached 2380 m/s (Mach number $M_{\mathrm{C}\mathrm{J}}=$ 6.74). In the CJ state, the temperature, pressure, and density of the detonation products were 3700 K, 2.94 MPa, and 2 kg/m³, respectively. Upon expansion to the baseline pressure ( $P_0=$ 0.1 MPa), the temperature of the detonation products was 2853 K (2580 °C). These thermodynamic equilibrium calculations were performed using the SDToolbox [53] and Cantera [54] software packages. When expanded to $P_0=$ 0.1 MPa and cooled to a characteristic thermal state of 2000 K, the evaluated composition of the detonation products comprised 65 vol.% H₂O, 32 vol.% CO₂, 2 vol.% CO, and 1 vol.% H₂. Thermodynamic modeling confirmed that this chemical composition remained frozen upon further cooling to lower temperatures.

3.3. Calculated Chemical Composition of Gasification Products

Thermodynamic simulations indicate that gasifying 1 kg of dry SM requires 0.3 kg of GA, establishing an optimal theoretical SM/GA mass ratio of ${\Omega }_{\mathrm{t}}\approx$ 3.3. At ${\Omega }_{\mathrm{t}}$ > 3.3 and gasification temperatures exceeding 1200 K, the solid carbon (graphite) yield increases up to 11 vol.%. This condensed-phase formation reduces the volume fractions of the primary off-gas components (H2 and CO), thereby lowering the total off-gas mass generated per kilogram of SM. Conversely, at ${\Omega }_{\mathrm{t}}$ < 3.3, the volumetric fractions of steam and carbon dioxide in the product stream increase, lowering the mass of dry off-gas and degrading the off-gas heating value due to the elevated CO₂ concentration. The calculated thermodynamic equilibrium compositions of the dry off-gas comprise 57 vol.% CO, 40 vol.% H₂, and 3 vol.% N₂ at temperatures above 1500 K. At a lower thermal state of 1000 K, the predicted profile shifts to 10 vol.% CO₂, 45 vol.% CO, 41 vol.% H₂, and 4 vol.% N₂, where nitrogen originates exclusively from the SM matrix.

3.4. Measurement Examples

Figure 3a,b illustrate representative transient profiles of the gas $T_{\mathrm{G}}$ and reactor wall $T_{\mathrm{W}}$ temperatures across two distinct experiments utilizing a 1-kg batch of raw SM (70% moisture) loaded via the upper hatch. In both trials, the PDG was activated at $t$ = 0. The thermochemical conversion of SM occurs in two sequential stages: moisture heating/evaporation and dry biomass gasification. The initial drying stage corresponds to the time interval ${\tau }_1$ , where $T_G$ establishes a distinct thermal plateau at 100 °C. The subsequent gasification stage is defined by the time interval ${\tau }_2$ , finishing when the diffusion burner flame is quenched. In the baseline run (Figure 3a), these time intervals were recorded as ${\tau }_1\approx$ 380 s and ${\tau }_2\approx$ 450 s, yielding a total process duration of ${\tau }_{\mathsf{\Sigma }}\approx$ 830 s. Upon completion, $T_{\mathrm{G}}$ remained 30–40 °C higher than $T_{\mathrm{W}}$ . Figure 3b details the corresponding real-time evolution of the off-gas composition (left scale). The gas analyzer response time varies by chemical species and is governed by sensor kinetics, the scrubbing media, and process homogeneity, with a minimum stabilization threshold of ~300 s. To account for these transient lags, replicate tests were performed to validate reproducibility. For the trial shown in Figure 3b, the drying and gasification stages lasted approximately 240 s and 600 s, respectively ( ${\tau }_{\mathsf{\Sigma }}\approx$ 840 s). The final dry off-gas comprised 32 vol.% CO₂, 40 vol.% CO, 21 vol.% H₂, 5 vol.% CH₄, and 2 vol.% C_xH_y (with 0 vol.% O₂), representing a total combustible fraction of 68 vol.%. Notably, the CO₂ concentration peaked during the first stage. This may be attributed to the initial influx of CO₂ from the GA stream before partial consumption via endothermic gasification reactions.

3.5. Measured Composition of Gasification Products

Table A4 in Appendix C compiles the parameters and outcomes of 17 representative experiments conducted on SM (runs #1 to #17). Gasification of the raw wet feedstock ( $\alpha =$ 70%) typically yielded a dry off-gas composed of 33–41 vol.% CO₂, 34–40 vol.% CO, 17–22 vol.% H₂, 2.5–4.0 vol.% CH₄, and 0–2.5 vol.% C_xH_y (Figure 4). Minimizing the CO₂ concentration directly correlated with an elevation in CO and H₂ fractions, whereas the hydrocarbon outputs remained largely invariant. For the pre-dried SM trials ( $\alpha =$ 15%), the CO₂ content dropped to 26–32 vol.%, while the peak concentrations of CO, H₂, and CH₄ rose to 45, 25, and 4.5 vol.%, respectively, pushing the total combustible fraction up to 74 vol.%. The lower heating value (LHV) of the effluent gas ranged between 7 and 12 MJ/kg. further enhanced by promoting the conversion of GA-derived CO₂ within the reactor. Parametric analysis indicated that under identical feedstock mass and moisture constraints, increasing the PDG pulse frequency accelerated the transformation rate, thereby shortening the total processing time (as demonstrated by comparing runs #2, #3, and #8 with run #4). Conversely, increases in either initial moisture content (runs #16 vs. #17) or batch mass (runs #10 vs. #17) clearly reduced the peak reactor wall temperature. These findings align with fundamental thermodynamic and heat transfer principles. The dataset establishes the isolated impacts of key operational boundaries, such as feeding configurations and initial wall temperatures, on the resulting off-gas composition, which are examined in the subsequent subsections.

3.6. Clarified Chemical Composition of the Off-Gas

In runs #8, #9, #10, and #11, gas sampling for chromatographic analysis via a Chromatec-Crystal 2000 gas chromatograph was conducted in parallel with real-time off-gas characterization using the MRU VARIO Plus flow-through analyzer. Table 1 compares the volumetric fractions of the gas components obtained by the flow-through analyzer against those from the gas chromatograph (indicated by asterisks, *). The chromatographic samples in runs #8 and #9 were collected downstream of the bubbler unit, whereas those in runs #10 and #11 were collected upstream of the bubbler. Comparative analysis showed that bubbling had a negligible effect on the final off-gas composition.

A comparison of the instrument outputs indicates that the volumetric fractions of CO and H₂ determined via gas chromatography deviate from the flow-through analyzer readings by an average of –3% (abs.) for CO and +5% (abs.) for H₂. Unlike the flow-through analyzer, which quantifies only methane and logs all other hydrocarbons (C_xH_y) as inert N₂, the gas chromatograph provides a breakdown of the C_xH_y fraction. Propane was identified as the predominant hydrocarbon detected, with a concentration below 0.1 vol.%. Beyond propane, the effluent off-gas contains trace amounts of ethylene and ethane, so the cumulative fraction of all detected hydrocarbons, including methane, remains below 5 vol.%. Consequently, the dry off-gas is free of hydrocarbons heavier than propane, confirming the absence of condensable organic fractions and tars.

3.7. Effect of the Feedstock Feeding Method and the Initial Temperature of the Reactor on the Dry Off-Gas Composition

As an alternative to the baseline batch configuration, a hydraulic piston feeder was implemented to deliver feedstock at a controllable, constant rate into the flow reactor cavity. This gradual feeding approach was expected to enhance mixing between the biomass and the GA, thereby improving carbon conversion efficiency, maximizing CO and H₂ outputs, and suppressing CO₂ formation. However, due to operational challenges stemming from the system’s transient pulsed backpressure, no noticeable improvements in gas quality have been achieved to date. Therefore, only preliminary findings for continuous processing are reported herein.

During these runs, the flow reactor was maintained at either a lower thermal state (200 °C) or an elevated temperature range of 400–500 °C. Figure 5 contrasts the resulting dry off-gas compositions obtained via batch loading (squares) and continuous feeding (diamonds) across various feedstock moisture contents (15–70 wt.%). The filled symbols represent a reactor wall temperature of 400–500 °C, whereas the empty symbols correspond to 200 °C. The data indicate that the continuous feeding configuration currently yields a slightly lower-quality gas profile than the baseline batch method, a limitation that is being addressed through ongoing feeder optimization. Regarding thermal conditioning, raising the initial reactor temperature consistently increased CO and H2 production while reducing CO₂ dilution, thereby improving the overall calorific value of the off-gas.

3.8. Effect of Feedstock Moisture on the Yield of Solid Residue

Experimental measurements indicate that the mass of solid residue recovered from the dry and wet cyclones ranges from 6.7% to 17.8% of the initial dry SM mass when processing raw wet feedstock ( $\alpha =$ 70%). Reducing the SM moisture content to $\alpha =$ 15% causes the solid residue mass to increase to 10–32.4%, demonstrating that pre-dried SM undergoes more intensive particulate entrainment from the flow reactor than the raw wet biomass, which aligns with expected fluidization and aerodynamic behaviors. To achieve complete conversion within the flow reactor, targeted measures must be implemented to mitigate this premature feedstock carryover. Evidently, the particle residence time inside the reaction zone must be extended, which can be accomplished by integrating intermediate buffer volumes between the PDG and the reactor or by deploying multi-stage cascading reactors.

3.9. Detonability of the Off-Gas

To characterize the detonability of the dry off-gas, an auxiliary gas accumulation and storage system was integrated into the experimental setup. This system included three wet coalescing filters with replaceable cartridges, a gas compressor with a 50-L tank and a peak capacity of 290 L/min, and a 200-L storage receiver. Pressure within the receiver was monitored via a pressure transducer rated up to 6.0 MPa. Off-gas accumulation experiments were conducted using SM with $\alpha =$ 15%. Composition analysis of the accumulated dry off-gas showed 32% CO₂, 43% CO, 21% H₂, 3% CH₄, and 1% C_xH_y. In the PDG, the stoichiometric mixture of off-gas and oxygen reached a mean detonation velocity of 1800 ± 100 m/s. Although this velocity falls below the measured detonation velocity for a stoichiometric NG–oxygen mixture (2100 ± 100 m/s) because of the lower heating value of the off-gas, its detonability remains adequate. Consequently, replacing NG with this off-gas can support a self-sustaining gasification process.

3.10. Elemental Composition and Oxidative Stability of the Solid Residue

The elemental composition of the solid residue recovered from the cyclone assembly (following drying) was evaluated for runs #8 and #9. In these trials, the GA mass flow rates were 1.82 g/s and 1.4 g/s, respectively, while the feedstock configurations were $M_0=$ 1000 g ( $\alpha =$ 70%) and 450 g ( $\alpha =$ 15%). The residual matrix for run #8 comprised 29.60 wt.% C, 7.50 wt.% O, 1.63 wt.% H, 1.64 wt.% N, and 59.63 wt.% ash, whereas run #9 yielded 42.13 wt.% C, 6.99 wt.% O, 1.98 wt.% H, 2.12 wt.% N, and 46.78 wt.% ash. These profiles demonstrate that the raw wet SM undergoes a deeper thermochemical conversion than the pre-dried feedstock. This behavior is evidenced by: (i) a lower residual carbon fraction in the wet SM char (30 wt.% vs. 42 wt.% for dried SM); (ii) an elevated ash concentration in the wet SM residue (60 wt.% vs. 47 wt.%); and (iii) a threefold reduction in the dry-basis carbon-to-ash ratio for the wet SM compared to the dried matrix. Consequently, the carbon conversion efficiency (CCE) reached 0.99 in run #8, compared to 0.67 in run #9. The near-complete CCE for raw SM is attributed to a gradual, layer-by-layer devolatilization mechanism. Conversely, the suppressed CCE for dried SM stems from rapid particulate entrainment driven by the high fluidizability and lower effective weight of dry particles under the shock-induced GA flow, leading to premature elutriation from the reactor core. Despite this, the lower heating value (LHV) of the off-gas derived from dried SM exceeded that of the wet SM feed. To reconcile these conflicting trends, implementing continuous feedstock injection instead of batchwise loading constitutes a critical technological requirement.

Figure 6a contrasts the mean mass fractions of various elements in the baseline dried SM with those in the solid residues recovered from the dry and wet cyclones (following condensate drying) across runs #1, #4, and #5. Elemental compositions were quantified via X-ray fluorescence (XRF) using an ARL PERFORM’X spectrometer (Thermo Fisher Scientific Inc., USA). As a representative example, minor variations in element concentrations between different runs are illustrated in Figure 6b with respect to phosphorus content. The elemental distributions within the dry residues of both cyclones were approximately identical. Given that the aggregate solid residue mass ( $m_s$ ) in these trials was a factor of 6–7 lower than the initial dry mass ( $M_d$ ) of the gasified SM batches, the absolute retention of specific elements varied. Notably, the relative concentrations of Ca, P, Al, and Mn in the dry SM were approximately 0.7, 0.5, 0.06, and 0.01 wt.%, respectively, but increased 6–7-fold in the solid residue to roughly 4, 3, 0.4, and 0.07 wt.%. For Si, Mg, Na, and Cl, their relative contents in the dry SM (1.2, 0.4, 0.2, and 0.1 wt.%, respectively) increased by only a factor of 2–3 in the solid residue, reaching 3.5, 1.2, 0.4, and 0.2 wt.%. The sulfur concentration remained nearly constant at approximately 0.3 wt.% in both the raw feedstock and the char. This disproportionate concentration effect suggests that elements such as Si, Mg, S, Na, and Cl partially volatilize into the gas phase, bypass the cyclone filtration train, or dissolve into the liquid condensate. The pronounced enrichment of Fe and depletion of V in the solid residue relative to the initial SM matrix warrant further investigation.

The oxidative stability of the solid residue was characterized by simultaneous thermal analysis (DSC/TGA) using an STA ZCT-1 apparatus (JING YI GAO KE, China). Char samples weighing 4.8 mg (sample #1) and 5.2 mg (sample #2) were placed in ceramic crucibles and heated in an air atmosphere from 50 to 950 °C at a constant heating rate of 10 °C/min. The thermokinetic profiles show a single prominent mass-loss step, quantified at 36.1% for sample #1 and 52.9% for sample #2. The derivative thermogravimetric (DTG) curves indicate that multiple overlapping degradation pathways of macromolecular segments occurred within the temperature ranges of 317–420 °C and 311–419 °C, respectively, accompanied by distinct caloric effects. The minor variations in the onset temperatures of these decomposition steps are attributed to chemical inhomogeneities between the analyzed samples.

3.11. Granulometric Analysis of the Solid Residue

Figure 7 contrasts the measured particle size distributions (PSDs) of the solid residues recovered from the dry and wet cyclones across runs #2, #5, #6, #8, and #9. The PSDs were quantified via wet laser diffraction using an Analysette 22 apparatus (Fritsch, Germany) in the presence of a surfactant under ultrasonic treatment (50 W). The solid particles exhibit diameters spanning 0.1 to 35 μm, with the PSD profiles remaining virtually invariant to changes in experimental conditions.

3.12. Chemical Composition of Liquid Condensate

Under the baseline batch loading configuration, where an initial period ( ${\tau }_1$ , Figure 3a) is dedicated to moisture vaporization, the feedstock simultaneously undergoes low-temperature pyrolytic decomposition promoting the formation of various intermediate fractions, including toxic pyrolytic by-products. To evaluate this phenomenon, a qualitative gas chromatography–mass spectrometry (GC-MS) analysis was performed on the liquid constituents recovered from the wet cyclone post-experiment. The analytical protocol and spectroscopic datasets are detailed in Appendix D. The primary identified compounds were phenol and coniferyl alcohol, while species such as toluene, propionic acid, furfuraldehyde, and furfuryl alcohol were detected in significantly lower concentrations. Following derivatization, the effluent was resolved as trimethylsilyl ethers, revealing an organic matrix rich in acids (propionic, butanoic, valeric, caproic, and benzoic), various furan derivatives, and benzyl alcohol. Thus, to suppress the formation of toxic condensable intermediates, the biomass must be injected continuously into a preheated reactor core. At specific GA mass flow rates and gas residence times, the optimized feedstock supply rate must guarantee the complete conversion of all carbonaceous fractions within the reactor.

4.1. Mass Balances

Data from runs #2, #3, #4, #8, and #12 enabled a precise determination of the mass ratios between the evaporated moisture, gasified dry matter, and supplied GA. Table 2 summarizes the duration of the feedstock moisture heating and evaporation stage ( ${\tau }_1)$ ; duration of the gasification stage of the dried feedstock ( ${\tau }_2$ ); total duration of the feedstock gasification process ( ${\tau }_{\mathsf{\Sigma }}$ ); initial feedstock batch mass ( $M_0$ ); moisture content ( $\alpha$ ); water mass ( $M_{\mathrm{w}}$ ); dry matter mass ( $M_{\mathrm{d}}$ ); alongside the mass flow rates of oxygen ( ${\dot{m}}_{\mathrm{o}\mathrm{x}}$ ) and NG ( ${\dot{m}}_{\mathrm{f}\mathrm{u}}$ ). Furthermore, the data provide the evaluated ratio of total consumed combustible mixture mass to initial feedstock batch mass ( $M_{\mathrm{f}\mathrm{o}\mathrm{x}}$ / $M_0$ ), where $M_{\mathrm{f}\mathrm{o}\mathrm{x}}=M_{\mathrm{o}\mathrm{x}}+M_{\mathrm{f}\mathrm{u}}$ ; mass of GA ( $M_{\mathrm{G}\mathrm{A}}$ ) consumed for gasification during the time interval ${\tau }_2$ ; mass of the produced dry off-gas ( $M_{\mathrm{g}}={(1-\beta )M}_{\mathrm{d}}+M_{\mathrm{G}\mathrm{A}}$ ), where $\beta$ is the ash content in the SM (8.58 wt.%); LHV of the dry off-gas ( $Q_{\mathrm{L}}$ , see Table A4); and the dry-basis cold gas efficiency (CGE). The CGE characterizes the process’s potential efficiency for downstream gas utilization. Assuming that high-temperature gasification reactions during the time interval ${\tau }_2$ occur exclusively between the GA and the dry feedstock, a dry-basis CGE estimate can be derived as follows [55]:

where $Q_{\mathrm{L},\mathrm{S}\mathrm{M}}$ corresponds to the LHV of the dry SM (14.6 MJ/kg) and $H_{\mathrm{G}\mathrm{A}}$ is the specific enthalpy of the GA calculated based on the GA composition at 2000 K (~3.6 MJ/kg).

As indicated by the final $M_{\mathrm{f}\mathrm{o}\mathrm{x}}$ / $M_0$ ratio in Table 2, gasifying 1 kg of wet SM during the time interval ${\tau }_{\mathsf{\Sigma }}$ requires about 1.52 kg and 1.35 kg of GA at moisture contents $\alpha =$ 70% and 45%, respectively. Linear extrapolation of these data to completely dry SM ( $\alpha =$ 0%) yields an experimental SM/GA mass ratio of ${\Omega }_{\mathrm{e}\mathrm{x}\mathrm{p}}\approx$ 1, implying that GA is consumed at a 1:1 mass ratio per unit of dry SM. Mitigating premature feedstock carryover enables generation of 1.91 kg of combustible off-gas, assuming a baseline ash content of approximately 9 wt.% in the dry SM. Consequently, the experimental feedstock-to-GA mass ratio was roughly 3.3 times lower than the optimized thermodynamic threshold, establishing a ratio of ${{\Omega }_{\mathrm{t}}/\Omega }_{\mathrm{e}\mathrm{x}\mathrm{p}}\approx$ 3.3.

The evaluated CGE values range from 121% to 158%, significantly exceeding the 100% threshold. Under the high-temperature conditions of the PDG, the external thermal energy of the GA is efficiently converted into the chemical energy of the resulting H₂- and CO-rich dry off-gas, thereby substantially elevating its cumulative heating value ( $M_{\mathrm{g}}{Q}_{\mathrm{L}}$ ) relative to the energy content of the dry feedstock ( $M_{\mathrm{d}}{Q}_{\mathrm{L},\mathrm{S}\mathrm{M}}$ ) and GA ( $M_{\mathrm{G}\mathrm{A}}{H}_{\mathrm{G}\mathrm{A}}$ ). CGE values exceeding 100% align with findings from several prior studies (e.g., [56], 124%).

To position the proposed technology among modern waste-to-energy pathways, Table 3 presents a formal comparison with mainstream biomass gasification technologies. The benchmark includes fluidized bed gasification, plasma gasification, and supercritical water gasification (SCWG), with a focus on key performance indicators: CGE, product quality, and economic feasibility, as highlighted in a recent review [57].

According to Table 3, while fluidized bed gasification is a commercially mature pathway, its application to complex feedstocks is often hindered by elevated tar yields and the need for syngas cleanup to remove toxic compounds and fly ash [58]. Unlike the complex remediation and carbon-recovery strategies required for conventional gasification residues, the proposed PDG technology inherently avoids forming such hazardous, complex carbon-ash structures and eliminates the need for extensive post-treatment. Similarly, plasma gasification generates high-quality, virtually tar-free syngas, yet its economic feasibility is compromised by intensive electricity consumption. SCWG is an effective route for wet biomass, but it requires high operating pressures, and its net energy efficiency is often compromised by the high thermal penalty needed to maintain supercritical water conditions.

In contrast, the proposed PDG technology offers a more balanced performance profile. Thus, in terms of the CGE, the results obtained in this work are considerably higher than those reported for dual bubbling fluidized bed biomass gasifiers (66% in [59]), values measured (24–51% in [60]) and estimated (up to 80% in [61]) for steam plasma-assisted gasification, as well as a value of 70% reported in [62] for SCWG. Moreover, the proposed system operates at ambient pressure and has significantly lower electrical consumption. Power demand is restricted to essential automation and utilities: the control system, a standard automotive spark plug for detonation ignition, and a minor consumption loop for the cooling water pump. It produces syngas of controllable quality under milder operating conditions, ensuring a competitive levelized cost of energy. In addition, the PDG system enables self-sustained operation by using its own syngas as fuel and offers straightforward scalability through adjustments to PDG and reactor dimensions, as well as PDG operating frequency. Nevertheless, potential explosion-safety constraints must be carefully managed. Further details regarding the economic feasibility and safety of the PDG technology are provided in [50].

4.2. Energy Balances

Runs #8 and #12 were evaluated to characterize the system energy balance. In run #8, a thermal input of 1.84 MJ was theoretically required to heat and vaporize the water within 1 kg of raw feedstock (70% moisture). Chronologically, during the initial drying stage ( ${\tau }_1=$ 380 s) under a PDG frequency of 0.63 Hz, the fuel consumption reached 114 g of NG. This corresponds to an energy input of 5.7 MJ, exceeding the theoretical enthalpy requirement by a factor of approximately 3. This discrepancy is attributed to parasitic, specifically heat dissipation into the reactor walls and the coolant in the PDG cooling jackets. Similarly, in run #12 ( $\alpha =$ 45%), containing 450 g of water with a theoretical vaporization enthalpy of 1.185 MJ, fuel consumption during the first stage ( ${\tau }_1=$ 260 s) was 72 g of NG. This translates to 3.57 MJ, representing a threefold oversupply. Importantly, both trials demonstrate a consistent relative thermal loss to the environment of approximately 67% during the initial stage, independent of the feedstock moisture content. Assuming equivalent heat losses govern the dry-SM conversion phase, it is inferred that only 33% of the GA thermal energy is directly utilized for the physical and chemical transformations of dry SM inside the flow reactor, while the remaining 67% is dissipated via wall conduction, jacket coolant loops, and sensible heat of the effluent off-gas. Incorporating this real thermal efficiency (33%) into the optimized theoretical feedstock-to-GA mass ratio ( ${\Omega }_{\mathrm{t}}\approx$ 3.3, as derived in the modeling section) yields a corrected practical estimate of ${\Omega }_{\mathrm{e}\mathrm{x}\mathrm{p}}\approx$ 0.33 ${\Omega }_{\mathrm{t}}\approx$ 1. This value agrees with the experimental mass balance results obtained via the extrapolation of wet SM data ( ${\Omega }_{\mathrm{e}\mathrm{x}\mathrm{p}}=$ 1.52 at $\alpha =$ 70% and ${\Omega }_{\mathrm{e}\mathrm{x}\mathrm{p}}=$ 1.35 at $\alpha =$ 45%) to a zero-moisture baseline.

4.3. Comparison of Experimental Data with Thermodynamic Calculations

When comparing the experimental datasets with the thermodynamic equilibrium predictions, several boundary constraints must be factored into the analysis. First, the experimental feedstock moisture ranged from $\alpha =$ 15% to 70%, whereas a zero-moisture baseline ( $\alpha =$ 0) was assumed in the calculations. Second, no thermal insulation measures were implemented on the laboratory setup, in contrast to the ideal adiabatic conditions modeled. Third, the measured detonation velocity within the PDG was lower than the theoretical ideal (2100 ± 100 m/s versus 2380 m/s), which altered the initial temperature and composition of the expanded detonation products at 0.1 MPa. Fourth, mixing between the GA and the condensed-phase biomass in the reactor required a finite induction time for particle heating and devolatilization, leading to spatial non-uniformities, whereas the model assumed instantaneous, molecular-level homogeneous gas-phase mixing. Finally, contrary to the infinite residence time assumed in the thermodynamic calculations, the real residence time was bounded, allowing for potential kinetic limitations and incomplete conversion.

The combined impacts of feedstock moisture and parasitic heat losses (the first two factors) on process thermal efficiency were quantified in the energy balance section, demonstrating that only 33% of the GA thermal energy directly drove the dry-SM phase and chemical transformations, while the remaining 67% was dissipated through wall conduction, cooling jackets, and sensible heat of the effluent off-gas.

Compared to these primary parameters, the influence of the reduced detonation velocity (the third factor) is relatively minor. A velocity deficit of approximately 12% ( $\left|2100-2380\right|\times 100/2380=$ 12%) corresponds to a GA temperature drop of less than 20% (from the ideal 2853 K down to ~2280 K), which remains sufficiently high to drive the target chemistry.

Table 4 highlights the structural consequences of finite mixing rates and spatial non-uniformities by comparing the predicted equilibrium and empirical compositions of the dry off-gas at 1000 K and 0.1 MPa. The final row lists the baseline mole fractions of CO₂, CO, and H₂ in the expanded detonation products exiting the flow reactor without biomass injection, assuming fully condensed water vapor.

Comparing the measured CO₂, CO, and H₂ concentrations in the effluent dry off-gas with their baseline values in the GA demonstrates that reactions yielding CO and H₂ proceed actively within the flow reactor, given that the initial volumetric fractions of these components in the GA were substantially lower. The equilibrium model generally predicts the CO fraction in the dry off-gas accurately. However, the experimental CO₂ content is 2–3 times higher, whereas the H2 output is 2–3 times lower than the thermodynamic predictions. Mechanistically, at temperatures exceeding 1200–1300 K, the generation of CO and H₂ is primarily governed by the following reactions:

where C denotes the carbonaceous matrix of the SM. Because the endothermic barrier of reaction (II) exceeds that of reaction (I), and the initial H₂O content in the GA is twofold higher than the CO₂ fraction (65 and 32 vol.%, respectively), the kinetic rate of reaction (II) can be significantly suppressed. This retardation is particularly pronounced at lower localized GA temperatures induced by parasitic heat sinks, specifically moisture vaporization and environmental dissipation. As a result, the input CO₂ component within the GA may undergo incomplete thermochemical conversion with the biomass, escaping the reactor vessel within the product stream. This mechanism accounts for the elevated residual CO₂ content observed in the dry off-gas at the terminal equilibrium state of 1000 K. The lower experimental hydrogen output is attributed to substantial spatial non-uniformities governing the heterogeneous reaction pathway (I).

To shift the actual off-gas composition toward the ideal thermodynamic equilibrium, future engineering efforts must focus on optimizing biomass–GA mixing, elevating the mean internal reactor temperature, and extending the hydrodynamics-driven particle residence time. No fundamental physical limitations hinder these operational modifications.

Experimental studies on the allothermal high-temperature gasification of swine manure were conducted across a range of feedstock moisture contents ( $\alpha =$ 15%, 45%, and 70%) using a custom laboratory-scale setup. The investigation yielded the following key conclusions:

• Off-Gas Quality and Biomass Drying: Gasification of raw wet SM ( $\alpha =$ 70%) produced a dry off-gas comprising 33–41 vol.% CO₂, 34–40 vol.% CO, 17–22 vol.% H₂, 2.5–4.0 vol.% CH₄, and 0–2.5 vol.% C_xH_y, achieving a carbon conversion efficiency of 0.99 and a cold gas efficiency ranging between 121% and 158%. Pre-drying the feedstock optimized the gas profiles, elevating the maximum yields of CO, H₂, and CH₄ to 45 vol.%, 25 vol.%, and 5 vol.%, respectively, while decreasing CO₂ to 26 vol.%. This modification increased the total combustible fraction up to 74 vol.%, although the CCE declined to 0.67 due to the premature elutriation of dry biomass particles.

• Solid Residue Properties: The particulate matter recovered from the cyclone system exhibited diameters confined to 0.1–35 μm, whereas the resultant particle size distributions demonstrated negligible sensitivity to variations in the operational parameters.

• Thermal and Mass Balances: In the current configuration, 67% of the input thermal energy from the high-temperature detonation products was dissipated via wall conduction, cooling jackets, and sensible off-gas heat, leaving a net thermal efficiency of 33% to drive the gasification chemistry. Under these constraints, the optimal practical feedstock-to-GA mass ratio was established as ${\Omega }_{\mathrm{e}\mathrm{x}\mathrm{p}}\approx$ 1, i.e., processing 1 kg of dry SM requires 1 kg of a stoichiometric NG—O₂ mixture, yielding 1.91 kg of combustible off-gas diluted with 26 vol.% CO₂.

• Model Validation and Scaling: The empirical datasets demonstrated parity with thermodynamic equilibrium simulations when factoring in structural heat losses. Moving the actual off-gas profiles toward ideal thermodynamic boundaries requires improving biomass–GA mixing, increasing the mean reaction-zone temperature, and extending particle residence time. No fundamental physical limitations hinder these system modifications.

Subsequent studies will evaluate the sustainability profile of the developed gasification technology through several key approaches, including reducing CO₂ concentrations in the off-gas, improving carbon conversion efficiency via continuous feedstock delivery into a preheated reactor core, maximizing the PDG pulse frequency to further accelerate transformation rates and minimize total processing time, and incorporating waste heat recovery systems for preliminary internal feedstock drying.

Conceptualization, writing—review and editing, resources, supervision, and project administration: S.M.F.; Methodology, writing—original draft preparation: S.M.F. and V.A.S.; Software: K.A.A.; Formal Analysis: S.M.F., V.A.S., I.A.S., and A.S.S.; Validation, A.S.S., I.A.S., and V.A.S.; Investigation, V.A.S., I.A.S., A.S.S.; Visualization, I.A.S., and V.A.S.; Data curation: F.S.F. and A.S.S. All authors have read and agreed to the published version of the manuscript.

The data will be available from the corresponding author on reasonable request.

The authors declare no conflicts of interest.

The study did not receive any external funding and was conducted using only institutional resources.

The authors are grateful to Aleksey B. Vorobyov, Aleksey V. Inozemtsev, Yaroslav O. Inozemtsev, Elena V. Koverzanova, Yurii A. Gordienko, Tatiana V. Dudareva, and Vera Ya. Popkova from the Semenov Federal Research Center for Chemical Physics, 4 Kosygin Str., 119991 Moscow, Russia, and Nikolai D. Blinov from the Federal Scientific Agroengineering Center VIM, 5, 1st Institutsky Proezd, 109428 Moscow, Russia, for their valuable assistance with measuring the physicochemical properties of the feedstock.

The authors confirm that no AI tools were used to generate any content of this manuscript. The authors take full responsibility for the accuracy, originality, integrity, and scientific content of the manuscript.

Appendix A.1. Gas Analysis

The volumetric composition of the generated off-gas was monitored using an MRU Syngas flow-through gas analyzer (Germany), which continuously recorded H₂, CO, CO₂, CH₄, O₂, and N₂ fractions with an estimated measurement error of 5%. Due to the analyzer’s lack of selectivity for non-methane hydrocarbons, the cumulative volume fraction of all C_xH_y compounds was inherently included within the N₂ balance. To improve conditioning and tar removal, the gas stream was routed through a bubbler filled with diesel fuel before analyzer injection. For more detailed effluent characterization, including individual C_xH_y species, discrete gas samples were simultaneously extracted for offline analysis using a Chromatec-Crystal 2000 gas chromatograph (Russia). To evaluate the scrubbing medium’s potential trapping or alteration effects on gas chemistry, sampling ports were positioned upstream and downstream of the diesel bubbler unit.

Appendix A.2. Pressure and Temperature Measurements

The upper section of the flow reactor was equipped with an overpressure sensor (Kurant DI, Russia; 1% accuracy) and two thermocouples configured to monitor the external reactor wall temperature (Pt100, Russia; accuracy $\pm$ 1 °C) and the local gas temperature (N-type DTPN285, Russia; accuracy $\pm$ 2.5 °C) near the internal wall boundary. The localized gas temperature reading was utilized strictly for stage identification rather than characterizing the core gasification temperature. The off-gas composition, internal reactor pressure, and upper-zone gas and wall temperatures were continuously recorded throughout each experimental run.

Appendix A.3. Analysis of Solid Residue

Following each experimental run, all solid residues were evacuated from the dual-cyclone system, with the fraction recovered from the wet cyclone subjected to a standard drying procedure. The resulting dry char was subsequently characterized using a comprehensive analytical suite: elemental CHNS analysis via a PE 2400 Series II automatic analyzer (Perkin Elmer, USA; 0.30% abs. accuracy); X-ray fluorescence (XRF) spectrometry on an ARL PERFORM’X instrument (Thermo Fisher Scientific Inc., USA); simultaneous thermal analysis (DSC/TGA) on an STA ZCT-1 apparatus (JING YI GAO KE, China); bomb calorimetry using an ABK-1V system (Russia); and particle size analysis via wet laser diffraction on an Analysette 22 apparatus (Fritsch, Germany) equipped with a 50-W ultrasonic dispersion unit and surfactant addition. The total mass and individual chemical compositions of the recovered cyclone residues were utilized to quantify the overall feedstock gasification efficiency.

Appendix A.4. Analysis of Liquid Condensate

The H₂O concentration in the produced off-gas was estimated from the mass of the condensate recovered from the wet cyclone. The chemical composition of this liquid phase was characterized via gas chromatography–mass spectrometry (GC-MS) using a Trace 1310 gas chromatograph coupled with an ISQ mass spectrometric detector (Thermo Fisher Scientific Inc., USA). To extract the organic constituents from the aqueous matrix, a liquid–liquid extraction protocol was performed by adding 4 mL of diethyl ether to a 4-mL condensate sample.

Table A1 summarizes the bomb calorimetry data for the higher heating value (HHV) of the SM, $Q_{\mathrm{H},\mathrm{S}\mathrm{M}}$ , obtained using an ABK-1V system (Russia). Because the raw feedstock has high moisture content, sample ignition required encapsulation in a 20 μm-thick terylene film pouch secured with a cotton thread. This assembly was placed in the calorimetric bomb crucible, with the thread connected to the ignition wire. The specific heat of combustion of the terylene film was specified as 22,790 J/g. Direct combustion of the SM sample without the auxiliary film was infeasible due to its elevated moisture content. In Table A1, $m_0$ , $m_{\mathrm{f}}$ , $m_{\mathrm{i}}$ , $m_{\mathrm{w}}$ , and $m_{\mathrm{s}}$ denote the masses of the feedstock sample, film, thread, water formed within the bomb crucible, and solid ash residue, respectively, while the higher heating value $Q_{\mathrm{H},\mathrm{S}\mathrm{M}}$ is calculated according to the following relation:

where $C$ denotes the heat capacity (energy equivalent) of the calorimeter, $∆T$ is the corrected temperature rise during the experiment, $q_{\mathrm{f}}$ is the combustion enthalpy of the film, $q_{\mathrm{i}}$ is the combustion enthalpy of the thread, and $q_{\mathrm{i}\mathrm{g}\mathrm{n}}$ is the electrical ignition energy. The mean higher heating value of the initial wet biomass was approximately 5.8 MJ/kg.

Table A2 presents the transient mass ( $m$ ) profiles of the SM samples as a function of drying time ( $t$ ) under ambient room temperature, where the final dry weight is designated as $m_{0\mathrm{d}}$ . The kinetic data indicate that mass loss ceases and stabilizes after approximately 48 h (two days) of exposure.

The dried SM sample was combusted in the calorimetric bomb without an auxiliary film. To increase the sample mass and enhance its representativeness, a specialized bomb configuration with two crucibles positioned at different heights within the cavity was utilized. The heat capacity (energy equivalent) of the calorimeter with this modified bomb assembly was calibrated in standalone measurements; it matched the single-crucible configuration once corrected for the individual heat capacities of the additional crucibles and fittings. The measurement data obtained for the air-dry SM specimens are compiled in Table A3 as runs #1 and #2. Supplementary vacuum drying of the air-dry sample (pressure $P\le$ 0.5 mm Hg) for 16 h induced a marginal mass reduction of approximately 0.3 wt.%. The combustion results for this vacuum-dried sample are designated as run #3 in Table A3, where the lower heating value ( $Q_{\mathrm{L},\mathrm{S}\mathrm{M}}$ ) is determined via the following expression:

Liquid condensate was characterized using a Trace 1310 gas chromatograph coupled with an ISQ mass spectrometric detector (Thermo Fisher Scientific, USA). Chromatographic separation was achieved on a TR-5MS fused-silica capillary column (15 m $\times$ 0.32 mm ID, 0.25 μm film thickness) with a 5% phenyl–95% methylpolysiloxane stationary phase. The oven temperature program started at 40 °C (held for 3 min), ramped to 120 °C at 5 °C/min, then to 300 °C at 15 °C/min, and finally held at 300 °C for 5 min. The injector temperature was maintained at 220 °C. Samples (1 μL) were injected using a split operating mode (split ratio 1:40). Mass spectra were acquired via electron ionization (EI) at 70 eV over a scanning range of 35–600 amu with a scan speed of 0.3 scan/s.

Liquid–liquid extraction was performed by adding 4 mL of diethyl ether to a 4-mL condensate aliquot. The mixture was continuously agitated for 10 min to ensure uniform mass transfer. After phase separation, the upper organic layer was collected and dried by filtration through a layer of anhydrous sodium sulfate on a paper filter. This extraction protocol was repeated twice using fresh solvent. The combined ether extracts were subsequently concentrated to a residual volume of 1–2 mL via rotary evaporation under vacuum at room temperature.

Silylation derivatization was performed using N,O-bis(trimethylsilyl)trifluoroacetamide (BSTFA) as the silylating agent. Initially, a 200-μL aliquot of the concentrated ether extract was evaporated to absolute dryness at 50 °C. The dry residue was treated with 20 μL of BSTFA and silylated at 50 °C for 15 min. The mixture was then reconstituted in 200 μL of hexane and analyzed via the aforementioned GC-MS protocol. Figure A1 illustrates the total ion current (TIC) chromatograms of the crude diethyl ether extract (Figure A1a) and the hexane extract containing the trimethylsilyl derivatives of the SM thermochemical conversion products (Figure A1b).

The following pyrolysis/gasification products were identified in the sample:

• -

  Main compounds (the most intense peaks on the chromatogram): phenol (emission time 7.49 min) and coniferyl alcohol (23.17 min);

• -

  Major compounds: toluene (1.77 min), propionic acid (1.96 min), furaldehyde (2.99 min), furfuryl alcohol (3.80 min), dimethylbenzene isomer (4.29 min), methyl phenol isomers (9.46 min and 10.14 min), dihydroxy benzene isomer (13.95 min), coumaran (14.23 min), syringol (17.58 min), vanillin (18.77 min), acetovanillone (20.46 min), syringyl aldehyde (22.48 min), dihydroxymethylguaiacol (22.58 min);

• -

  Minor compounds: ethylphenol isomers (12.71 min and 12.78 min), methylguaiacol (13.20 min), hydroxybenzaldehyde isomer (18.29 min), acetosyringol (23.12 min).

Compounds characteristic of lignin pyrolysis products was identified in the SM pyrolysis/gasification products: methylguaiacol, syringol, acetovanillone, vanillin, and others. It can be assumed that the SM pyrolysis/gasification products also contain other compounds typical of lignin pyrolysis products: various acids and alcohols, which were not detected because the stationary phase of the TR-5MS capillary column used in GC-MS analysis was weakly polar and was not suitable for analyzing such highly polar compounds as alcohols and acids, especially linear and low-molecular-weight ones.

After derivatization, the following products were identified as trimethylsilyl ethers (Figure A1b):

• -

  Acids: C5 oxyacid with a molecular weight of 116 (2.15 min), propionic acid (2.34 min), methylpropionic acid (3.04 min), butanoic acid (4.33 min), methylbutanoic acid isomers (4.76 min and 5.71 min), valeric acid (7.08 min), C6 acid with a molecular weight of 116 (7.83 min), caproic acid (9.92 min), benzoic acid (14.71 min), hydrocinnamic acid (19.17 min);

• -

  Furan derivatives: furfuryl alcohol (7.62 min), coumaran (14.18 min). Phenols: phenol (9.16 min), methylphenol isomer (11.39 min, 11.75 min, and 12.03 min), dihydroxybenzene monoester isomer (13.44 min), guaiacol (14.22 min), ethylphenol isomer (14.49 min), dihydroxybenzene diester isomer (16.70 min), hydroxybenzaldehyde (18.00 min), dimethoxyphenol (18.88 min), hydroxyacetophenone (20.16 min), vanillin (21.11 min);

• -

  Alcohols: benzyl alcohol (12.14 min).

Analysis of the silyl derivatives also showed that the pyrolysis/gasification products of SM are similar to the pyrolysis products of lignin.