# Biomass gasification in an autothermal semi-industrial fluidized bed gasifier: Syngas characterization and energy balance

P O Badu, M Debal, P Girods, S Aubert, Y Rogaume

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P O Badu, M Debal, P Girods, S Aubert, Y Rogaume. Biomass gasification in an autothermal semi-industrial fluidized bed gasifier: Syngas characterization and energy balance. Biomass and Bioenergy, 2025, 197, pp.107831. 10.1016/j.biombioe.2025.107831. hal-05202799

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# Biomass gasification in an autothermal semi-industrial fluidized bed gasifier: Syngas characterization and energy balance

P.O. Badu ID, M. Debal, P. Girods, S. Aubert, Y. Rogaume*

Université de Lorraine, INRAE, LERMAB ERBE, Épinal, F - 88 000, France

## ARTICLE INFO

### Keywords:

- Gasification

- Bubbling fluidized bed

- Biomass

- Cold gas efficiency

- Energy balance

## ABSTRACT

Biomass gasification is well-established in industrial applications, but semi-industrial autothermal systems remain underexplored despite their role as a bridge between laboratory scale and full-scale deployment. Understanding syngas characteristics, energy recovery, and system efficiency at this scale is critical for optimizing industrial gasification. This study examines air gasification of woody biomass in a 200 kW autothermal bubbling fluidized bed gasifier operating at equivalence ratios (ER) 0.22-0.45 with recorded bed temperatures of 695-860 °C. Key performance indicators, including syngas composition, lower heating values (LHV), cold gas efficiency (CGE), carbon conversion efficiency (CCE), and energy distribution of the near-industrial system, were analyzed. Results show syngas primarily contained H₂ (6.4-11.3 %), CO (8.5-16.6 %), CH₄ (3.0-5.5 %), and CO₂ (15.6-19.4 %), with LHV of 3.9-7.1 MJ/Nm³. CGE reached 51.8-71.4 %, and CCE 78.6-98.5 %. Overall system efficiency (79-93 %) highlights the role of energy recovery from syngas cooling and gathered solid residues in improving large-scale gasification viability. This study provides novel insights into optimizing semi-industrial gasification, demonstrating the feasibility of integrated energy recovery for enhanced efficiency and sustainability.

## 1. Introduction

Global energy demand is growing exponentially due to rapid economic development and a gradual population increase. Environmental concerns have resulted in the reassessment of energy sources, necessitating a progressive shift away from fossil fuels. Emissions from fossil fuel combustion threaten human health and the ecosystem. Limited nature of fossil fuel reserves is an added reason for the energetic transition to more sustainable and renewable energy sources, providing long-term energy security with minimal environmental impacts. Biomass is highly regarded as one of the suitable renewable alternatives to fossil fuels with a huge potential [1].

Biomass energy is a significant renewable source because of its sustainable carbon cycle. It accounts for nearly 14 % of the world's primary energy, potentially reaching 40 % by 2050 [2]. It also presents positive environmental and socioeconomic benefits accompanied by innovative and conventional approaches to generating transportation fuels, heat, and power [2]. However, the potential overuse of biomass resources remains a key challenge. Biomass production inherently depends on land, creating inevitable competition between land use for food

production and industrial applications. Consequently, assessing biomass availability for industrial purposes must consider these limitations. A balanced approach involving a combination of agricultural residues, waste streams, and purpose-grown energy crops is recommended. Estimates suggest that up to 1 billion tonnes of biomass could be sourced, with approximately 50 % derived from agricultural residues and waste and 50 % from energy crops specifically cultivated for this purpose [3]. Biomass is generally valorized for energetic applications through three thermochemical conversion approaches: direct combustion, pyrolysis, and gasification, with gasification being the most efficient among these three for small to medium-scale electricity production [4].

Over the past two decades, significant advancements have been made in gasifier designs for converting biomass into syngas. Despite these efforts, only a limited number of large-scale systems (>100 MWe) have successfully transitioned from demonstration to achieve sustained commercial operation. In contrast, small-to-medium scale systems have demonstrated greater operational success [5-7]. For instance, the Billerud Karlsborg mill in Sweden successfully implemented biomass gasifiers for electricity generation during the BIGDME (Biomass Integrated Gasification Dimethyl Ether) and BIGCC (Biomass Integrated Combined

<footnote>* Corresponding author.</footnote>

<footnote>E-mail address: yann.rogaume@univ-lorraine.fr (Y. Rogaume).</footnote>
Cycle) projects [8]. Similarly, a gasifier unit (48 MW) constructed by Carbona/Andritz at a Finnish pulp mill in Joutseno has replaced 95 % of the natural gas consumption in its lime kiln. Andritz has further advanced gasification technology by constructing a 65 MW unit at Zhanjiang Chenming Pulp & Paper Mill in China, successfully replacing heavy oil in its kiln. Additionally, an 80 MW gasifier plant built by Andritz for the Chenming Group is operational at the Shouguang Meilun greenfield pulp mill in China [9]. These examples highlight the growing adoption of biomass gasification technologies in small-to-medium-scale industrial applications, indicating their role in enhancing sustainability and reducing dependence on fossil fuels.

Gasification is a commercialized thermochemical technology that applies high temperatures (>650 °C) under controlled oxidizing atmospheres to convert carbonaceous solid fuels such as biomass or waste into intermediate gases. Combustible gases produced from gasification contain primarily H₂, CO, CH₄, CO₂, and N₂ [10,11]. The syngas also contain particulates, tar, ammonia, and hydrogen sulfide in minute quantities depending on the nature of the biomass feedstock. These compounds, tar in particular, cause clogging and corrosion issues during downstream application hence an appropriate cleaning process is required for their removal [10–12]. Cleaned syngas can be used for chemical synthesis, and burnt directly as fuel in boilers or gas engines [12]. Although biomass gasification is mainly used for combined heat and power generation, the process may also produce energy vectors such as methane or hydrogen. Many investigations have been conducted over different scales to study and develop these technologies [13,14].

Depending on the syngas application, different oxidizing agents involving air, oxygen, oxygen-enriched air, steam, and oxygen-steam mixtures can be used for gasification. While gasification with oxygen-steam mixtures yields hydrogen-rich syngas with a superior heating value, gasification in air produces a non-reactive nitrogen-diluted combustible gas with an inferior heating value [14]. However, for optimum operational and maintenance cost, air was used as an oxidizing agent in this study, avoiding the complex process of oxygen separation or steam generation that comes at an additional cost. Biomass gasification research is broadly classified into two main paths: direct and indirect. With the direct process, the reactor is operated as autothermal, that is, part of the fuel is oxidized to generate the heat required to drive endothermic reactions during gasification. Indirect gasification, on the contrary, is allo-thermal, meaning an external heat source is needed to drive endothermic gasification reactions [15,16].

Biomass gasification and performance indicators involving syngas yield and quality, cold gas, and carbon conversion efficiency are influenced by the choice of biomass feedstock, oxidizing agents, gasification operating conditions, and reactor type [17]. Extensive research has been conducted on lignocellulosic biomass gasification in fluidized bed reactors using air as an oxidizing agent. Karatas et al. [12] conducted indirect air and steam gasification on a cotton stalk and hazelnut shells in an atmospheric bubbling fluidized bed reactor. In their work, the equivalence ratio (ER) which is defined as the actual air-fuel ratio divided by the stoichiometric air-fuel ratio, determined by the elemental composition of each feedstock, was varied from 0.36 to 0.71 for cotton stalks and 0.25 to 0.68 for hazelnut shells. The steam-to-biomass ratio was increased from 0.52 to 1.69 and 0.33 to 2.26 for cotton stalk and hazelnut shells, respectively. Equivalence ratio (ER) deeply influenced product syngas, yielding lower heating values (LHV) within the ranges of 2.49–3.24 MJ/Nm³ and 3.05–7.83 MJ/Nm³ for cotton stalk and hazelnut shells, respectively.

Mohammed et al. [18] also studied air gasification of empty fruit bunch for hydrogen-rich syngas production using an indirectly heated bench-scale fluidized bed reactor. They varied reactor temperature from 700 to 1000 °C, and equivalence ratio (ER) in a range of 0.15–0.35. Results from their study show high reactor temperature favors hydrogen-rich syngas formation. With an optimum equivalence ratio (ER) of 0.25 and a reactor temperature of 850 °C, hydrogen-rich syngas (H₂ - 27.31 %vol) was achieved. In another work, Pio et al. [17]

performed direct (air) gasification of wood pellets, pine, and eucalyptus residual forest biomasses (RFB) in an auto-thermal bubbling fluidized bed. They varied the equivalence ratio from 0.17 to 0.36 and gasification temperatures between 700 and 800 °C. Their results indicate syngas with superior LHV are better formed at a low equivalence ratio.

Karmakar et al. [19] also investigated the influence of equivalence ratio (ER) (0.25–0.45) and bed temperature (600–800 °C) on fuel gas production using an air-blown bubbling fluidized bed gasifier on rice husk. They also observed decreased syngas quality at a higher equivalence ratio (ER). This observation was attributed to a decrease in the H₂ and CO concentrations at elevated ER resulting in N₂ diluted syngas with higher CO₂ concentration. However, at a fixed ER, H₂ and CO concentrations increased whereas CH₄ and CO₂ decreased at higher reactor temperatures.

From the literature [12,17–19], it is clear that gasification conditions influence greatly syngas quality. LHV of syngas generally increases with decreasing equivalence ratio while increasing reactor temperature at a fixed equivalence ratio favors a hydrogen-rich syngas formation. Downstream syngas application is, therefore, a major deciding factor in the choice of biomass feedstock, gasification parameters, and gasifier type. Majority of existing literature on biomass gasification with bubbling fluidized beds is based on laboratory-scale gasifiers [12,18, 20–27], however, pilot-scale gasifiers are gaining grounds lately [10,11, 17,19,28]. Pilot-scale atmospheric bubbling fluidized beds are one of the simplest and most economically viable gasifiers for biomass gasification. They are reliable with many feedstocks at varying particle sizes and have a robust technology for the power range between 1 and 10 MWel. They also exhibit a high and uniform heat transfer between the bed material and biomass [29,30]. Previous studies on lignocellulosic biomass gasification in pilot-scale bubbling fluidized bed gasifiers focused primarily on gasification conditions, and syngas characteristics to evaluate gasification performance [10,11,17,19]. However, there is a noticeable lack of studies addressing the energy analysis of lignocellulosic biomass gasification on a semi-industrial scale. This gap limits the understanding of energy efficiency and recoverability in near-commercial operations, critical for scaling up to industrial applications.

To bridge this research gap, the authors conducted an energy analysis of multiple commercial woody biomasses with distinct properties, utilizing a 200 kW atmospheric bubbling fluidized bed gasifier. This study evaluated energy distribution within the pilot plant during gasification and quantified the recoverable energy during syngas cleaning and cooling processes. The findings have significant implications for optimizing energy utilization in downstream applications, enhancing the overall performance of industrial-scale gasifiers. The study provides novel insights into energy efficiency and distribution across various biomass types, offering a practical understanding of gasification dynamics under semi-industrial conditions. Specifically, it examines the gasifier performance during syngas cooling using heat exchangers and a Venturi scrubber, presenting a detailed energy recovery analysis. By comparing results with existing literature, this work highlights key differences in gasification behavior between the tested biomasses, shedding light on the interplay between biomass characteristics, gasification conditions, and performance metrics such as cold gas efficiency and carbon conversion. Syngas composition, a critical indicator of gasification success, is also presented, with an analysis of how varying operational conditions influence its quality. The study's findings bridge the gap between laboratory-scale experiments and real-world industrial applications, offering actionable insights for project development and process optimization.

A comprehensive description of the experimental facility and methodology is provided in Section 2, ensuring reproducibility and transparency. Section 3 delves into detailed observations and results, accompanied by a thorough discussion of their implications for industrial-scale biomass gasification. Key findings and broader implications of this work are summarized in Section 4, underscoring its contribution to advancing sustainable energy technologies and guiding
future industrial applications.

## 2. Materials and methods

### 2.1. Gasification facility

The gasification facility used in this study is a semi-industrial scale pilot plant designed specifically for advanced research applications, provided by EQTEC. This facility features an atmospheric bubbling fluidized bed gasifier capable of processing biomass feedstocks at a flow rate of 50 kg/h with a nominal output power of 200 kW. Although the plant is categorized as a pilot-scale unit, it integrates all the key components found in full-scale commercial gasification systems, enabling the production of high-quality, useable syngas suitable for various downstream applications.

The gasifier is engineered to accommodate a wide range of biomass feedstocks with differing physical characteristics, including particle sizes from 10 mm to 50 mm and densities between 150 kg/m³ to 700 kg/m³. This flexibility allows for experimental runs using various biomass types, such as pellets and wood chips, making it an ideal platform for studying the effects of feedstock variability on gasification performance. Designed to simulate industrial conditions, the facility enables precise control over operational parameters and supports detailed energy and mass balance analysis. A schematic layout of the pilot plant is depicted in Fig. 1. This setup provides a valuable basis for scaling research findings to larger, commercial applications and enhances obtained data relevance for real-world implementation.

The feedstock is stored in a skip with drying capabilities and is initially fed into hopper 1 (1), where nitrogen blanketing ensures an inert environment, preventing unintended oxidation. Feedstock is transferred into hopper 2 (2) to safeguard against uncontrolled air entry, a crucial step for stable reactor operation. Fuel is then sent into the reactor via two sequentially connected screws. The first screw (3) acts as a dosing mechanism, calibrated for each test run, with its rotational speed finely tuned to achieve target flow rates. This calibration involves temporarily disassembling screws 3 and 4 and using a load cell at screw 3's outlet to directly measure fuel mass, establishing a precise, linear relationship between fuel mass flow and screw speed.

The second screw (4) transfers biomass into the reactor (7) while maintaining a controlled disconnection from the dosing screw (3), which mitigates pyrolysis front progression and ensures operational safety. Biomass injection occurs at 0.86 m above the distribution plate. A rotary lobe compressor (5) facilitates air injection, while a natural gas burner (6) preheats this air during the start-up phase. The reactor (7)

itself, with a height of 5 m, consists of two sections: a lower section (3 m high with an internal diameter of 0.45 m) and an upper section (2 m high and an internal diameter of 0.85 m). Its walls are insulated with 0.1 m of refractory concrete and 0.1 m of insulation board to minimize thermal loss.

The syngas cleaning system is optimized for effective residue and tar removal. A cyclone separator (8) with a cooling hopper collects fly ash, coarse particulates, and other residues. While the cyclone primarily removes dry fly ash, minor quantities of fly ash-tar agglomerates may also be present due to early stage tar condensation on ash surfaces at elevated temperatures (~550 °C); The cyclone separator unit has a 0.36 m internal diameter, a 0.2 m exhaust diameter, a 0.992 m body height, and a 1.2 m conical height, with wall composition matching the reactors. Connected via a U-shaped insulated pipe, the air heat exchanger (AHE) (9) cools syngas to 300 °C, using 24 pipes (0.027 m diameter, 3.4 m height) to facilitate syngas flow and simultaneously provide heated air for bed fluidization. A baghouse filter (10) with nine e-glass fabric bags (1.8 m high, 0.12 m in diameter) further cleanses the syngas by capturing fine particulates (including residual fly ash, fly ash-tar agglomerates, and unconverted carbon particles). The syngas is cooled again to 50 °C by a water heat exchanger (WHE1) (11), and subsequently treated in a venturi scrubber (12) (0.0224 m diameter) and a spray tower (13). At this final stage of tar removal, fine water droplets facilitate the separation of residual light tars and tar aerosols that persist in the syngas after prior condensation in the upstream heat exchangers and filtration units. While a significant fraction of heavy tars condense at higher temperatures (~300 °C), lighter tars such as benzene and toluene remain gaseous until exposed to deep scrubbing action of the Venturi scrubber. However, while this process significantly reduces tar content, complete tar elimination is not achievable, as trace amounts of ultrafine tars may remain in the cleaned gas. A final water heat exchanger (14) cools syngas to 15 °C, condensing residual water.

To ensure operational safety, a rotary lobe compressor (15) maintains a slight negative pressure (–2 mbar) in the reactor to prevent syngas leakage, essential as the pilot operates indoors. Cleaned syngas is then combusted in a boiler (16) with a premix burner that supports syngas-natural gas co-combustion for transitional phases. Water used for syngas cleaning undergoes treatment in a dissolved air flotation unit (DAF) (17), allowing for recycling within the system. Excess water containing residual tars is collected and stored before disposal. This pilot setup is technically robust and representative of full-scale systems, providing a reliable foundation for upscaling research outcomes to industrial applications.

<figure><img src="image_3.png" /><figcaption>Fig. 1. Schematic representation of the gasification facility.</figcaption></figure>
## 2.2. Measurements

The pilot plant has a comprehensive array of sensors to monitor gasification conditions and conduct mass balance analysis for each experiment. Airflow rate is precisely measured using an orifice meter and a hot wire sensor, ensuring reliability of these critical measurements, which are essential for determining equivalence ratio (ER), syngas production flow, and overall mass balance. The reactor has 15 sample points, 14 temperature probes, and 12 pressure probes distributed along its height. The syngas cleaning line features a sampling point, a temperature probe, and a pressure probe between each component. All sensors and pilot parameters are recorded every second during experimental runs.

Gas composition was measured at the process outlet, following the rotary lobe compressor (15), using three dedicated sampling lines. The first sampling line maintained raw gas at 250 °C and diluted it 15 times with N₂ for online tar analysis, specifically targeting benzene, and toluene, using a µ-gas chromatography (µGC SRA R 3000) equipped with a StabiliWAX DB column. The µ-gas chromatography was calibrated with a custom-prepared gas cylinder to obtain three calibration points 500, 1000, and 2000 ppm for benzene and two calibration points 250 and 1000 ppm for toluene. Off-line gas chromatography with a flame ionization detector (GC-FID) was used to confirm and validate the benzene concentration previously measured. The GC-FID was calibrated by injecting a commercially available standard solution, which is seriously diluted to generate multiple calibration points (1000, 500, 200, 100, 50, and 10 µg/mL). The resulting calibration curves for both compounds exhibited perfect linearity within these ranges.

In the second sampling line, the gas was cooled via three condensers to remove water and condensable gases, then filtered through dry sawdust and wool to eliminate residual particulates. Cleaned gas was analyzed online using a µ-gas chromatography (SRA R 3000) equipped with PoraplotU and Molsieve 5A columns to determine the concentrations of CO, CO₂, CH₄, C₂H₂, C₂H₄, C₂H₆, H₂, N₂, and O₂. To verify these measurements, a second µ-gas chromatography with the same configuration was used in series with the first, ensuring accurate quantification of the permanent gases. The third sampling line was employed for offline tar measurement. The raw gas was transferred through a heated line (maintained at 250 °C) to a tar sampling system, following the standard tar protocol [31]. This system consisted of four impingers in series (the first at 0 °C, the remaining three at -20 °C), each containing 60 ml of isopropanol. The gas flow in this line was monitored after the tar sampling flow rate was approximately 1 L/min, with each tar measurement lasting 6 min. The collected solvent was analyzed using gas chromatography with a flame ionization detector (GC-FID) (Shimadzu GC 2010 Plus), equipped with a Zebron ZB-1MS column (Phenomenex) for aromatic compound separation.

## 2.3. Experimental procedure

Each experimental run on this pilot installation spans an average of five days. Due to the reactor's semi-industrial scale, substantial 5-ton refractory concrete, and complex design, it requires a minimum of 40 h to achieve a steady state. Experiments are conducted in three key phases: preheating, combustion, and gasification. Detailed start-up procedure for the installation is outlined below.

Starting at room temperature, a bed of 150 kg of olivine, approximately 0.6 m high, is fluidized in a bubbling regime using preheated air at 450 °C. Fluidization air used to preheat the reactor is achieved via a natural gas burner (6) between the air blower and plenum. After 12 h, once the bed reaches 380 °C, feedstock is introduced into the reactor at a low flow rate to enable complete combustion. Heat generated from the complete combustion of feedstock sustains the reactor temperature. The preheater is subsequently deactivated. The bed then continues to heat and stabilizes at target gasification temperature after an additional 24 h. During this period, feedstock and airflow rates are incrementally

adjusted to reach set values for the experiment. A further 12 h of heating, all components in the system reach operational temperatures, resulting in steady-state conditions with average recorded temperatures of 600 °C in the reactor freeboard, 550 °C in the cyclone separator (8), and 450 °C at the first heat exchanger (9) inlet. Testing commences after a total of 48 h of heating to steady-state conditions. Between each experimental condition, a stabilization period of at least 2 h is maintained for reliable data collection. This systematic process allows the gasifier to sustain autothermal operation by balancing heat generation from feedstock combustion with heat consumption by endothermic reactions in the system.

Test parameters, including feedstock type, air, feedstock flow rates, and their corresponding equivalence ratios (ER), are summarized in Table 1. In an autothermal reactor, ER and gasifier bed temperature are interdependent; thus, isolating the effect of ER at a constant temperature as typically done in indirectly heated gasifiers, [10,12,20,32,33], is not feasible. Consequently, the temperatures reported in Table 1 reflect interdependent variables, influenced by feedstock rate and ER. The semi-industrial scale and autothermal operation of this reactor yield results that closely approximate conditions found in full-scale industrial gasification setups. However, given this reactor's relatively higher surface area-to-volume ratio, it exhibits greater heat loss per unit surface area than fully industrial reactors, which benefit from enhanced thermal insulation.

After the gasifier unit was shut down, residues separated by the cyclone separator and bag filter were gathered and weighed. Ash content, moisture content, and higher heating value (HHV) were measured. The particle production yield is the average of an operating period, including the start-up, combustion, warm-up, and shut-down phases. It

<table><caption>**Table 1**
Gasification conditions in the semi-industrial bubbling fluidized bed gasifier.</caption><thead><tr><th>Biomass Reference</th><th>Fuel Type</th><th>ER</th><th>Recorded Gasifier Bed Temperature (°C)</th><th>Biomass feed rate (kg/h) d.b</th><th>Air flowrate (Nm³/h)</th></tr></thead><tbody><tr><td>P</td><td>Wood pellets</td><td>0.36</td><td>860</td><td>38</td><td>62</td></tr><tr><td>P</td><td>Wood pellets</td><td>0.33</td><td>859</td><td>41</td><td>62</td></tr><tr><td>P</td><td>Wood pellets</td><td>0.25</td><td>832</td><td>55</td><td>62</td></tr><tr><td>P</td><td>Wood pellets</td><td>0.26</td><td>817</td><td>52</td><td>62</td></tr><tr><td>P</td><td>Wood pellets</td><td>0.22</td><td>792</td><td>56</td><td>56</td></tr><tr><td>C1</td><td>Woodchips</td><td>0.35</td><td>706</td><td>43</td><td>73</td></tr><tr><td>C1</td><td>Woodchips</td><td>0.30</td><td>695</td><td>50</td><td>73</td></tr><tr><td>C1</td><td>Woodchips</td><td>0.36</td><td>701</td><td>48</td><td>84</td></tr><tr><td>C1</td><td>Woodchips</td><td>0.34</td><td>775</td><td>48</td><td>79</td></tr><tr><td>C2-A</td><td>Woodchips</td><td>0.34</td><td>769</td><td>61</td><td>100</td></tr><tr><td>C2-A</td><td>Woodchips</td><td>0.35</td><td>774</td><td>52</td><td>86</td></tr><tr><td>C2-A</td><td>Woodchips</td><td>0.35</td><td>772</td><td>64</td><td>106</td></tr><tr><td>C2-A</td><td>Woodchips</td><td>0.36</td><td>787</td><td>61</td><td>107</td></tr><tr><td>C2-A</td><td>Woodchips</td><td>0.37</td><td>793</td><td>57</td><td>100</td></tr><tr><td>C2-A</td><td>Woodchips</td><td>0.38</td><td>793</td><td>55</td><td>100</td></tr><tr><td>C2-A</td><td>Woodchips</td><td>0.38</td><td>788</td><td>50</td><td>92</td></tr><tr><td>C2-A</td><td>Woodchips</td><td>0.42</td><td>800</td><td>49</td><td>100</td></tr><tr><td>C2-A</td><td>Woodchips</td><td>0.43</td><td>795</td><td>48</td><td>100</td></tr><tr><td>C2-B</td><td>Woodchips</td><td>0.38</td><td>740</td><td>47</td><td>86</td></tr><tr><td>C2-B</td><td>Woodchips</td><td>0.39</td><td>746</td><td>46</td><td>87</td></tr><tr><td>C2-B</td><td>Woodchips</td><td>0.39</td><td>733</td><td>46</td><td>86</td></tr><tr><td>C2-B</td><td>Woodchips</td><td>0.39</td><td>776</td><td>46</td><td>87</td></tr><tr><td>C2-B</td><td>Woodchips</td><td>0.40</td><td>745</td><td>45</td><td>87</td></tr><tr><td>C2-B</td><td>Woodchips</td><td>0.42</td><td>745</td><td>43</td><td>87</td></tr><tr><td>C2-B</td><td>Woodchips</td><td>0.44</td><td>789</td><td>47</td><td>100</td></tr><tr><td>C2-B</td><td>Woodchips</td><td>0.45</td><td>793</td><td>46</td><td>100</td></tr></tbody></table>

P – wood pellet C1 – woodchip with 23.1 % moisture C2 – woodchip with 29.2 % moisture.

C2-A – woodchip (C2) with high feeding rate C2-B – woodchip (C2) with low feeding rate d.b – dry basis.
was estimated as the ratio of gathered residues from the cyclone separator and bag filter to the quantity of biomass injected in the reactor.

## 2.4. Fuels

Three commercially sourced woody biomasses, designated P, C1, and C2 were used as fuels for gasification experiments. Biomass P consisted of wood pellets (30 mm in length and 6 mm in diameter) suitable for small boilers and stoves, characterized by low moisture content. In contrast, Biomass C1 and C2 were two distinct wood chips for industrial boiler use featuring a maximum particle size of 50 mm and higher moisture content. Results of the elemental analysis (following the ISO 29541 standard method), moisture content (determined by oven drying by ISO 11722), ash content (measured by combustion in a muffle furnace at 815 °C according to ISO 18122), and calculated lower heating value (LHV) of the fuels are detailed in Table 2.

LHV was calculated considering moisture content, H content, and HHV. HHV was measured using a Parr 6100 calorimeter. Using these results, the stoichiometric air volume for the combustion of 1 kg of biomass was calculated.

## 3. Results and discussion

Results obtained from the gasification experiments provide critical insights into the influence of key process parameters on syngas composition, tar yield, and overall system performance. The study highlights the role of fuel properties, particularly moisture content, in determining gasification efficiency, product gas composition, and energy recovery. Temperature distribution within the gasifier was monitored, revealing distinct thermal zones that directly impacted reaction pathways and syngas quality. The influence of equivalence ratio (ER) was particularly significant, affecting combustion intensity, tar cracking efficiency, and the balance between oxidation and reduction reactions. Higher ER values enhanced oxidation, increasing reactor temperatures and promoting secondary cracking of hydrocarbons, leading to improved syngas quality with reduced tar content. However, excessive ER values also led to increased CO₂ formation, which may decrease syngas heating value. Additionally, several performance parameters were calculated to assess the gasification efficiency. Gas yield was determined to evaluate the volume of syngas produced per unit of biomass input, while heating values of the generated gas were analyzed to determine its energy potential. Cold gas efficiency (CGE) was assessed as a key metric of process performance, providing insights into the fraction of the biomass chemical energy converted into syngas. Carbon conversion efficiency was quantified, offering an indication of the extent to which carbon from the feedstock was utilized in gas production. An energy balance analysis was also conducted at selected operational points, providing a comprehensive evaluation of heat distribution and energy recovery within the system. Subsequent sections explore detailed effects of these variables, beginning with the role of equivalence ratio.

### 3.1. Influence of equivalence ratio

Equivalence ratio (ER) was varied between 0.22 and 0.45, yielding recorded bed temperatures within the range of 695–860 °C. As no

<table><caption>**Table 2**

Elemental analysis, moisture content, and LHV of fuels.</caption><thead><tr><th></th><th>LHV (MJ/ kg)</th><th>Moisture content (wt %, w. b.)</th><th>Ash (wt%, d. b.)</th><th>C (wt %, d. b.)</th><th>H (wt %, d. b.)</th><th>O wt %, d. b.)</th><th>N (wt %, d. b.)</th></tr></thead><tbody><tr><td>P</td><td>16.95</td><td>7.9</td><td>0.3</td><td>49.6</td><td>6.6</td><td>43.4</td><td>&lt;0.03</td></tr><tr><td>C1</td><td>13.49</td><td>23.1</td><td>1.0</td><td>50.6</td><td>6.5</td><td>41.8</td><td>0.1</td></tr><tr><td>C2</td><td>12.38</td><td>29.2</td><td>1.0</td><td>50.6</td><td>6.5</td><td>41.8</td><td>0.1</td></tr></tbody></table>

d.b: dry basis w.b: wet basis.

external heating source was applied during gasification, feedstock, and airflow rates were carefully modulated to achieve steady-state conditions. Consequently, bed temperature was sustained solely through heat generated by exothermic reactions, specifically carbon oxidation, char oxidation to carbon monoxide, the water-gas shift, and char hydrogenation respectively (reactions 1–4) [18,34,35].

$$C + O_2 \rightarrow CO_2 - 393.8 \text{ kJ/mol} \quad (1)$$

$$C + 0.5O_2 \rightarrow CO_2 - 110.5 \text{ kJ/mol} \quad (2)$$

$$CO + H_2O \rightleftharpoons H_2 + CO_2 - 41 \text{ kJ/mol} \quad (3)$$

$$C + 2H_2 \rightleftharpoons CH_4 - 74.9 \text{ Kj/mol} \quad (4)$$

As ER increased, additional oxygen was introduced to the gasifier, thereby intensifying exothermic oxidation reactions involving carbon in the biomass and char. This, in turn, elevated bed temperatures, as illustrated in Fig. 2. For wood pellets (P), an increase in ER from 0.22 to 0.36 raised the bed temperature from 790 to 860 °C. Similar trends were observed for wood chips (C1 and C2), demonstrating consistency in thermal behavior across different feedstocks. Notably, an unexpected temperature spike occurred for wood chips (C1) at an ER of 0.34. While this anomaly remains unexplained, it highlights the complexity and sensitivity of gasification thermodynamics, potentially pointing to unaccounted factors such as local bed heterogeneity or rapid char reactivity.

For wood chips (C2), results were stratified into two distinct groups based on biomass feeding rates: C2-A (48–64 kg/h, dry basis) and C2-B (43–47 kg/h, dry basis). Within each group, the ER-to-temperature relationship as illustrated in Fig. 2, was consistent with observations for P and C1, supporting similar findings in the literature [17,18,36]. Grouping was essential, as higher feed rates at similar ERs generally result in elevated bed temperatures due to increased fuel power input into the reactor [36]. This separation allowed for more precise data interpretation across a range of operating conditions.

A key finding was the requirement of a higher ER for wood chip feedstocks (C1 and C2) compared to wood pellets (P) to attain comparable temperatures. This observation is attributable to the higher moisture content in wood chips (23.1 % - 29.1 %) relative to pellets (7.9 %). Elevated moisture necessitates additional energy for evaporation, consequently reducing bed temperatures for wood chips compared to pellets at equivalent ERs [37–39]. These findings emphasize the critical role of feedstock moisture in thermal regulation during gasification and underscore the importance of adjusting ER according to biomass characteristics to optimize gasifier performance.

<figure><img src="image_4.png" /><figcaption>**Fig. 2.** Influence of equivalence ratio (ER) on gasifier bed temperature for each stabilized period.</figcaption></figure>
Due to the maximum allowable gasifier bed temperature, experimental runs at higher ER values where bed temperatures exceeded 860 °C were not conducted. Similarly, experiments at ER values below 0.2 were infeasible as limited exothermic reaction activity made it impossible to maintain bed temperature without external heating. Within the permissible range, ER values varied between 0.22 and 0.36 for wood pellets (P), 0.30 and 0.36 for woodchips C1, and 0.34 and 0.45 for woodchips C2.

Fig. 3 illustrates longitudinal temperature profile along the reactor length at similar equivalence ratios for all three biomass types (wood pellets: 0.36, woodchips C1 and C2: 0.35). A consistent thermal behavior was observed across all three feedstocks, characterized by a gradual temperature decline from the bed region (0.66 m above the air distributor plate) to the reactor exit. The highest temperatures were recorded in the bed zone due to oxygen availability, which facilitated partial and complete oxidation of char, leading to significant thermal energy release from the biomass. Beyond the bed zone, reactor temperature exhibited a decreasing trend. In the biomass injection zone, an initial temperature drop was observed, likely attributed to the endothermic nature of biomass drying and devolatilization process [40]. These reactions require energy input to drive feedstock conversion into volatile gases and solid char, temporarily reducing local temperatures before exothermic oxidation reactions dominate downstream. As the reaction zone transitions towards the freeboard region near the gasifier exit, a continuous decline in temperature is observed. Oxygen depletion progressively suppresses exothermic oxidation reactions, resulting in reduced thermal energy generation [17]. Endothermic secondary tar cracking reactions further influence the temperature gradient, as heavier hydrocarbons undergo thermal decomposition into lighter gas species, absorbing heat in the process. Heat dissipation through the reactor walls also plays a critical role in the observed temperature profile, exacerbating thermal losses along the freeboard region [41]. The interacting thermal phenomena highlights the intricate heat transfer dynamics within the gasifier and emphasizes the necessity for precise control of operating conditions and reactor insulation for optimum process efficiency.

The study revealed a significant influence of ER on syngas composition across all biomass types. The primary gaseous components N₂, CO₂, CO, H₂, CH₄, and C₂H₄ were detected in decreasing order of concentration, with trace amounts of C₂H₂, C₂H₆, C₆H₆, and C₇H₈ also present. Specifically, N₂ concentrations ranged from 48 to 55 %vol for wood pellets and 55–62 %vol for woodchips while CO₂, CO, H₂, CH₄, and C₂H₄ varied between 15.6 and 19.4 %vol, 8.5–16.6 %vol, 6.4–11.3 %vol, 3.0–5.5 %vol, and 0.8–1.7 %vol respectively. These results align well with existing results on syngas profiles from woody biomass gasification [17]. As shown in Fig. 4, N₂ concentration is significantly higher for

<figure><img src="image_5.png" /><figcaption>Fig. 3. Temperature distribution profile along reactor length.</figcaption></figure>

woodchips than wood pellets. This is attributed to the lower equivalence ratio (ER) used for wood pellets, which results in reduced airflow and, consequently, less nitrogen dilution. The influence of ER on nitrogen content is evident across all biomass types. The higher N₂ content in syngas from woodchips as shown in Fig. 4, is primarily due to the higher ER required for these feedstocks relative to wood pellets. Higher ER involves increased airflow, resulting in greater nitrogen dilution. This trend, which corresponds to the reduced concentrations of combustible gases, has also been noted in previous studies, including those by Motta et al. [37].

To accurately quantify ER impact on syngas composition, recalculations of gas concentrations were conducted on an N₂-free basis, eliminating nitrogen's effect (Fig. 5). The recalculated values reveal CO₂ concentrations increase with ER across all biomass types. Higher oxygen availability promotes oxidation of both biomass and char, leading to elevated CO₂ formation. For wood pellets, CO₂ concentration peaked at an ER of 0.33 with a minimal decrease at an ER of 0.36. This reduction in CO₂ correlates with modest increases in CO and H₂ concentrations, likely a result of higher gasifier temperatures shifting equilibrium toward CO and H₂ formation in favor of endothermic reactions [42]. The behavior underscores critical interplay between ER and gasifier temperature in controlling product gas composition. Higher gasifier temperatures at a constant ER promote reforming reactions, which favor combustible gases (CO and H₂) formation at the expense of CO₂.

Although woodchips C1 (23.1 % moisture) and C2 (29.2 % moisture) exhibited similar elemental compositions, their differing moisture content had a measurable impact on dry syngas composition. As shown in Table 1, at the same ER of 0.35, woodchip C2 produced slightly higher H₂ concentrations (7.7 % and 8.8 %) compared to woodchip C1 (7.6 %). This increase in H₂ yield can be attributed to the elevated moisture content in C2, which enhanced methane reforming reactions, promoting CH₄ conversion into H₂. Conversely, CO₂ levels in syngas from woodchip C2 (17.8 % and 19.0 %) were marginally lower than those of woodchip C1 (19.3 %) at the same ER. This reduction can be attributed to increased CO₂ transformation into CO at slightly elevated bed temperatures of 774 °C, and 772 °C for woodchip C2 compared to the lower 695 °C recorded for woodchip C1. These findings indicate minor variations in feedstock moisture content can influence syngas composition by affecting the extent of reforming and shifting reactions. This relationship is consistent with the results of Karmakar et al. [19], who observed lower CO₂ concentrations at elevated temperatures due to enhanced activity of endothermic reforming reactions such as the Water-gas, Boudouard, and methane reforming reactions outlined below [19,34,43]:

$$C + H_2O \rightleftharpoons CO + H_2 + 131.4 \text{ kJ/mol} \quad (5)$$

$$CH_4 + H_2O \rightleftharpoons CO + 3H_2 + 206 \text{ Kj/mol} \quad (6)$$

$$C + CO_2 \rightleftharpoons 2CO + 172 \text{ kJ/mol} \quad (7)$$

This behavior aligns with Le Chatelier's principle, whereby an increase in temperature favors endothermic reactions, enhancing CO and H₂ production. In contrast, for exothermic reactions like carbon oxidation, a rise in temperature shifts the equilibrium back toward reactants, minimizing CO₂ formation [29].

For woodchips (C1), CO₂ concentration initially dropped at an ER = 0.34 before rising with increasing ER, accompanied by a reduction in combustible gases. High bed temperature at ER = 0.34 (775 °C) likely promoted reforming reactions, enhancing H₂, CH₄, and C₂H₄ production. A similar trend was observed for woodchips C2, where higher ERs led to increased CO₂ and decreased combustible gas concentrations, aligning with findings from related studies [19,44]. Overall, these findings reinforce the established relationship between ER and syngas composition of woody biomass gasification. Higher ERs promote oxidation and CO₂ formation but reduce energy-rich gases, highlighting the need for ER optimization to maximize syngas quality, especially in
<figure><img src="image_6.png" /></figure>

<figure><img src="image_7.png" /></figure>

<figure><img src="image_8.png" /><figcaption>Fig. 4. Effect of Equivalence Ratio (ER) on principal dry gas composition (N₂, H₂, CH₄, CO, CO₂, C₂H₄) for a) wood pellet, b) woodchip (C1), and c) woodchip (C2) fuels.</figcaption></figure>

authothermal systems where temperature control depends on ER and reaction balance.

The Lower Heating Value (LHV) of syngas was calculated based on its composition, including benzene and toluene on a dry basis using the following equation:

$$LHV_{\text{syngas}} = \sum X(i) \cdot LHV_{(i)} \quad (8)$$

Where X(i) denotes the volume fraction of each gaseous species (i = CO, CH₄, CO₂, C₂H₄, C₂H₂, C₂H₆, C₆H₆, and C₇H₈) in the syngas and LHV(i) represents the lower heating value of species (i) in MJ/N m³.

The effects of ER on syngas LHV are represented in Fig. 6. Across all three biomass types, an inverse linear relationship was observed between ER and syngas LHV. At higher ERs, increased oxygen availability enhances carbon combustion resulting in syngas with high nitrogen and CO₂ concentrations. This nitrogen dilution and reduced formation of combustible gases, such as CH₄, CO, and C₂H₄, which are primary contributors to syngas LHV, lead to lower energetic content in the syngas [18,29]. This effect explains why syngas produced from woodchips exhibited a lower LHV compared to that from wood pellets. Furthermore, the reduced LHV at higher ERs correlates with a decrease in tar content, specifically benzene and toluene, as shown in Fig. 7. Together, due to higher bed temperature and oxygen availability, both of which favor tar-cracking reactions [45,46]. Although benzene and toluene represent over 80% of total tar mass, and their concentrations decrease with increasing ER, benzene and toluene are present in low concentrations (<1%) in dry syngas, their heating values contribute 3–9% to the

overall syngas LHV. The higher benzene content observed in syngas from wood pellets (0.24–0.38 %vol) compared to woodchips (0.09–0.18 %vol) further enhances the LHV of syngas derived from wood pellets. At a similar ER around 0.35 for all three biomass types, the LHV of syngas from wood pellets was significantly higher than that from woodchips (C1 and C2). This difference can be partially attributed to the higher bed temperature achieved during wood pellet gasification, promoting carbon conversion and forming higher-LHV combustible gases. Conversely, lower bed temperatures during woodchip produced syngas with higher nitrogen and CO₂ concentrations, as previously shown in Fig. 4, resulting in reduced syngas quality and LHV. The maximum observed LHV (7.1 MJ/Nm³) was achieved during wood pellet gasification at an ER of 0.22, while the lowest LHV (3.9 MJ/Nm³) was observed during woodchip (C2) gasification at an ER of 0.45. Fig. 6 illustrates the effects of the ER on the LHV of syngas highlighting the relationship between ER and syngas energy content.

Tar content in the syngas is also significantly influenced by gasification parameters, particularly equivalence ratio (ER) and bed temperature. Data from the pilot plant indicates increasing ER values lead to a noticeable reduction in tar yield, demonstrating the effectiveness of autothermal tar cracking mechanisms. This trend is consistent across all fuel types (wood pellets and woodchips, C1, and C2). As the ER increased from 0.22 to 0.36 for wood pellets, tar yield decreased from 5.4 g/Nm³ to 2.4 g/Nm³, correlating with a temperature rise from 792 °C to 860 °C. This suggests higher ER values enhance oxidation reactions, increasing thermal energy and promoting secondary tar cracking. A similar trend is observed for woodchips C1, where increasing
<figure><img src="image_9.png" /><figcaption>Fig. 5. Influence of Equivalence Ratio (ER) on dry gas composition, calculated on an N₂ free basis, for a) wood pellet, b) woodchip (C1), and c) woodchip (C2) fuels.</figcaption></figure>

<figure><img src="image_10.png" /><figcaption>Fig. 6. Impact of equivalence ratio on lower heating value of dry syngas.</figcaption></figure>

the ER from 0.30 to 0.34 resulted in tar yield reduction from 3.0 g/Nm³ to 2.3 g/Nm³, with temperatures rising from 695 °C to 775 °C. However, at ER = 0.36, tar yield slightly increased to 3.2 g/Nm³, indicating that beyond a certain threshold, additional oxidation may promote char combustion rather than further tar cracking. For woodchips C2 with higher moisture content (29.2 %), the tar yield follows a progressive

decline as ER increases. At ER = 0.34, tar yield is 3.5 g/Nm³, decreasing to 2.8 g/Nm³ at ER = 0.39, with temperatures ranging from 769 °C to 793 °C. The slightly higher initial tar yield compared to C1 can be attributed to moisture-enhanced steam reforming reactions which influence tar breakdown differently than direct thermal cracking. These findings highlight the impact of moisture content on gasification behavior, with moisture content initially yielding more tar but benefiting from steam-assisted reforming, leading to tar reduction at elevated temperatures. The influence of gasification parameters on the tar yield is illustrated in Fig. 8.

Syngas production yield quantifies the volume of dry gas (including nitrogen) generated per kilogram of biomass fed into the reactor. Yield was calculated using nitrogen from the air oxidizing agent as an internal standard according to Eqn (9).

$$ \text{Gas yield} = \frac{Q_{\text{air}} \cdot N_2, \text{air}}{Q_{\text{fuel}} N_2, \text{syngas}} \qquad (9) $$

Where $Q_{\text{air}}$ represents the air volumetric flowrate (Nm³/h), $Q_{\text{fuel}}$ represents biomass feeding rate (kg/h), $N_2, \text{air}$ represents the percentage volume of N₂ in the air (79 %vol), and $N_2, \text{syngas}$ represents the percentage volume of N₂ in dry syngas (%vol). While syngas LHV decreased with increasing ER, an appreciable increase in gas yield was observed. This increase in yield can be attributed to higher oxygen availability and increased bed temperature at elevated ERs, which enhance tar cracking, char oxidation, and the combustion of combustible gases, producing more syngas per unit biomass, as shown in Fig. 9. At higher ERs, the ratio of fuel to airflow decreases, which lowers the overall volumetric flow of gas at the reactor exit due to reduced biomass input. However, when
<figure><img src="image_11.png" /></figure>

<figure><img src="image_12.png" /><figcaption>Fig. 7. Equivalence ratio effect on the concentrations of a) benzene and b) toluene in the dry syngas.</figcaption></figure>

<figure><img src="image_13.png" /><figcaption>Fig. 8. Trends in tar yield under varying Equivalence Ratio and gasifier bed temperature.</figcaption></figure>

normalized per kilogram of biomass fed into the reactor, gas yield increases with rising ER. This trend aligns with the behavior of syngas obtained with air gasification, which is primarily nitrogen-dominated, as observed in previous studies [18,29]. Dry syngas yield ranged from 1.5 to 2.1 $Nm^3/kg_{biomass}$, consistent with the results reported by Tim-sina et al. [47].

Examining the three types of biomass (P, C1, and C2) reveals feed-stock characteristics play a significant role in gas yield. Wood pellets demonstrated the highest yield at an ER of 0.36, whereas the lowest yield was recorded at an ER of 0.22. Across similar ER values, wood pellets consistently produced higher yields than woodchips. This is likely due to the higher moisture content in woodchips, which reduces bed temperature and, consequently, gas production efficiency. Similar trends were reported by Lv et al. [48] who observed a decrease in gas yield (2.53–1.43 $Nm^3/kg_{biomass}$) with decreasing temperature from 900 to 700 °C during sawdust gasification.

## 3.2. Cold gas efficiency and carbon conversion

Cold gas efficiency (CGE) is another key metric for assessing gasification performance, defined as the ratio of the power in the produced syngas to the combustible power of the biomass feedstock, expressed as

<figure><img src="image_14.png" /><figcaption>Fig. 9. Impact of equivalence ratio on the yield of dry syngas produced.</figcaption></figure>

a percentage. CGE was calculated using the estimated gas yield and syngas LHV (from section 3.1) according to Eqn. (10):

$$ \text{CGE} = \frac{\text{Gas yield} \cdot \text{LHV}_{\text{syngas}} * 100}{\text{E}_{\text{biomass}}} \quad (10) $$

Where $E_{biomass}$ is the fuel power from the biomass (kW). The CGE values observed for all experimental runs ranged from 51.8 % to 71.4 %, with the highest efficiency achieved during wood pellet gasification at an ER of 0.36 and the lowest during woodchip C1 gasification at 0.30. These CGE values fall within the upper range reported for air gasification in bubbling fluidized beds [17]. Across all feedstocks, CGE increased with rising ER, largely due to higher gas yields at elevated ERs. Notably, wood pellet gasification consistently resulted in higher CGE than woodchip gasification, despite the latter requiring a higher ER range. This indicates that factors such as bed temperature and biomass properties also significantly impact CGE. The high moisture content in woodchips, for instance, likely lowered the bed temperature during gasification, which in turn limited gas yield and consequently reduced CGE.

In this study, CGEs above 50 % were achieved at ERs between 0.22 and 0.45, aligning with the range reported by Pio et al. [17]. Similarly,
Karmakar et al. [19] found CGEs greater than 50 % at ERs of 0.25 and 0.35 when bed temperatures exceeded 650 °C and 700 °C, respectively. Thus, bed temperatures above 650 °C combined with ERs in the range of 0.2–0.4 appear optimal for achieving high CGE in woody biomass gasification. ER and recorded bed temperature influence on CGE illustrated in Fig. 10, demonstrates increasing ER and temperature positively impact CGE by enhancing syngas production efficiency.

Carbon Conversion Efficiency (CCE) represents the fraction of carbon in the biomass feedstock converted into gaseous products. CCE was calculated using Eqn. (11) [18].

$$CCE (\%) = \frac{\sum_i \delta c_i \cdot i \times C(i) \times Mc \times Q_{gas}}{M_v \times Q_{biomass} \times C_{f,db}} \times 100 \quad (11)$$

Where i denotes the primary carbon-containing compounds in syngas (CO, CO₂, CH₄, C₂H₄, C₂H₂, C₂H₆, C₆H₆, and C₇H₈), δc, i is the number of carbon atoms present in each compound (δc, CO, CO₂, CH₄ = 1, δc, C₂H₄, C₂H₂, C₂H₆ = 2, δc, C₆H₆ = 6, δc, C₇H₈ = 7), C(i) represents the volume fraction of each carbon-containing compounds in the dry syngas, M_c is the molar mass of carbon (g/mol), Qgas denotes the volumetric gas flow (Nm³/h), M_v is the molar volume of gas at normal conditions (22.4 × 10⁻³ Nm³/mol), Qbiomass is the biomass feed rate (kg/h), and C_f.db is the fraction of carbon in biomass on dry ash free basis.

For all experimental runs, CCE improved with increasing ER and recorded bed temperatures, which can be attributed to enhanced reactivity of biomass at elevated temperatures and increased oxidizing agent availability at higher ERs. These conditions facilitate greater carbon conversion into syngas components. Calculated CCE values ranged from 79 % to 98.5 %, aligning well with values reported in the literature [18]. The maximum carbon conversion efficiency was achieved during woodchips (C2) gasification at an equivalence ratio of 0.45, while the minimum was observed for woodchips (C1) at an ER of 0.30. Unlike cold gas efficiency, which was notably higher for wood pellet gasification, similar CCE ranges were observed across all biomass types. However, achieving comparable CCE values for woodchips (C1 and C2) required higher ERs due to their higher moisture content, which necessitated greater oxidizing conditions for effective carbon conversion, as illustrated in Fig. 11.

### 3.3. Pilot efficiency

The pilot system's efficiency was evaluated by analysing energy distribution throughout the gasification process, focusing on syngas production, heat recovery from syngas cleaning, cooling, and solid residues. Heat recovery rates from the system's components were expressed

<figure><img src="image_15.png" /><figcaption>Fig. 11. Effect of equivalence ratio on carbon conversion efficiency in biomass gasification.</figcaption></figure>

in kilowatts (kW) to represent instantaneous thermal power transfer, providing a dynamic assessment of energy flows. Specifically, thermal energy recovered rate during syngas cooling in the Air Heat Exchanger (AHE) to 300 °C (E_AHE), and in the Water Heat Exchanger (WHE1) from 220 °C to 50 °C (E_WHE1) were calculated using Equations (12) and (13), respectively. Gas temperature upstream of the AHE was estimated based on the bed temperature, yielding a range of 530–650 °C, while all other temperatures were measured during steady-state operations. In WHE1, additional energy was recovered from the enthalpy of water vaporization. A hydrogen balance was performed for each test period to determine the water content in the syngas to calculate the wet syngas enthalpy. The syngas heat capacity was calculated according to its temperature using values from the literature [49]:

$$E_{AHE} = \dot{V}_{syngas} \cdot C_{PAHE} \cdot \Delta T_{AHE} \quad (12)$$

$$E_{WHE1} = \dot{V}_{syngas} \cdot C_{WHE1} \cdot \Delta T_{WHE1} + m_{H_2O} \cdot \Delta vap H \quad (13)$$

Where Vsyngas is the syngas volumetric flow rate (Nm³/h), CPAHE and CPWHE1 are the estimated syngas volumetric heat capacities during cooling in the air and water heat exchangers respectively (kJ/Nm³°C); ΔT_AHE

<figure><img src="image_16.png" /><figcaption>Fig. 10. Effect of a) equivalence ratio and b) gasifier bed temperature on Cold gas efficiency (CGE) of wet woody biomass in a bubbling fluidized bed.</figcaption></figure>
and $\Delta T_{WHE1}$ denote the temperature drops during syngas cooling in the respective exchangers, $m_{H_2O}$ represents the mass flowrate of water present in the syngas (kg/h), and $\Delta_{vapH}$ represents the enthalpy of water vaporization (2270 kJ/kg).

The recoverable thermal energy rate from syngas cooling in the AHE ranged from 13.5 to 26.1 kWth, with the lowest value observed for woodchip C1 and the highest for woodchip C2. Energy rate from WHE1 was within a lower range of 13.6–19.0 kWth for wood pellets, increased to 24–27.8 kWth for woodchip C1, and reached an upper range of 28.2–40.2 kWth for woodchip C2. This variation correlates with the moisture content and enthalpy of the syngas produced during woodchip gasification.

In addition to syngas, the energy balance considered solid residues (fly ash, fly ash-tar agglomerates, and any remains of partially unconverted carbon particles) collected in the cyclone and bag filter from start-up to shutdown phase. The total particle yield (residues gathered from cyclone and bag filter) ranged from 2.4 to 5 % of the biomass input. These residues had a higher heating value (HHV) of approximately 25 MJ/kg, representing 5–9 % of the fuel energy input to the gasifier. In an industrial setting, such residues could be used as fuel. The energy recoverable rate from the solid residues was calculated as follows:

$$E_{cyc} = m_{cyc} \cdot HHV_{cyc} \qquad (14)$$

$$E_{BF} = m_{BF} \cdot HHV_{BF} \qquad (15)$$

Where $E_{cyc}$ represents the energy recoverable rate from solid residues in the cyclone (kW), $m_{cyc}$ is the mass flow rate of solids in the cyclone (kg/h), $HHV_{cyc}$ is the higher heating value of solid residue recovered from the cyclone (kJ/kg), $E_{BF}$, $m_{BF}$, $HHV_{BF}$ represent the energy recoverable rate, mass flow rate, and higher heating values of solid residues in the bag filter, respectively.

The overall gasification efficiency of the pilot plant was calculated as follows:

$$\eta_{overall} = \text{CGE} + \frac{E_{cyc} + E_{AHE} + E_{BF} + E_{WHE1} * 100}{E_{biomass}} \qquad (16)$$

Where $\eta_{overall}$ denotes overall gasification efficiency, while $E_{cyc}$, $E_{AHE}$, $E_{BF}$, $E_{WHE1}$, and $E_{biomass}$ represent the recoverable energy rates from the cyclone, air heat exchanger, bag filter, water heat exchanger, and fuel power in the biomass respectively, as calculated in previous sections. Overall gasification efficiencies, illustrated in Fig. 12, show a range of 79 %–93 % across all experimental runs. The minimum efficiency was observed for woodchip C1 at a bed temperature of 695 °C, while the maximum was achieved for wood pellets at 860 °C. Despite moderate

<figure><img src="image_17.png" /><figcaption>Fig. 12. Influence of gasifier bed temperature on overall gasification efficiency of wet woody biomass.</figcaption></figure>

CGE values for wet biomass, the overall gasification efficiencies were consistently above 70 %, underscoring the importance of energy recovery from both syngas cooling and solid residues in optimizing gasification performance.

Similar to cold gas efficiency (CGE), overall gasification efficiency increases with rising bed temperature for all biomass types. However, within comparable temperature ranges, wood pellets exhibited lower overall gasification efficiency than woodchips. This difference is primarily due to the lower instantaneous thermal power transfer from heat exchangers in the syngas cleaning line during wood pellet gasification compared to woodchip gasification. The reduced heat recovery results in lower overall gasification efficiency for wood pellets relative to woodchips at equivalent bed temperatures, as shown in Equation (16). Conversely, overall gasification efficiency for biomass with consistent moisture content improves as the reactor bed temperature increases. This study demonstrates high overall gasification efficiency can be achieved with wet biomass, although a balance between moisture content and overall gasification efficiency is essential, as excessive moisture impacts gasification performance.

To detail accumulated energy distribution within the pilot plant during gasification and syngas cleaning, a comprehensive energy balance was conducted at six selected test points. Localization of energy concentration areas within the gasifier unit was categorized into four main areas: (1) recoverable energy from residues, (2) heat exchangers, (3) syngas, and (4) thermal loss from the reactor (1 - $\eta_{overall}$). Results from the energy balance at the selected test points highlighting the energy concentration regions on the pilot plant are shown in Fig. 13. The analysis confirms that moisture content in the biomass significantly influences the amount of heat recovered from the exchangers. For each biomass type, heat recovery from exchangers increased with rising ER due to the corresponding increase in reactor bed temperature during gasification. The energy balance results show minimal difference in residue yield regardless of biomass moisture content. Moreover, the estimated thermal losses decreased consistently with increasing ER. Specifically, for wood pellets, thermal loss decreased by a factor of 2.7 as ER increased from 0.22 to 0.36. In the case of C1 woodchips, the reduction in thermal loss was by 1.5 when ER increased from 0.30 to 0.34. Similarly for C2 fuel, a reduction factor of 1.7 was observed as ER rose from 0.40 to 0.45. These findings underscore the critical role of ER in optimizing thermal efficiency within the gasification process, with increasing ER values contributing to a more favorable energy balance.

The observed reduction in thermal losses, calculated as the difference between the overall gasification efficiency and unity (1 - $\eta_{overall}$), can be attributed to the increased energy content of the syngas and enhanced recovery of sensible heat from the heat exchangers as ER rises. As previously discussed, a higher ER improves carbon conversion efficiency, allowing a greater proportion of the feedstock’s chemical energy to be transformed into syngas, thereby improving overall gasification efficiency. This increased conversion efficiency reduces the amount of thermal energy lost to the environment, resulting in a marked decrease in thermal losses. In essence, the conversion of a substantial portion of the feedstock’s chemical energy into syngas, coupled with optimal recovery of heat from the process maximizes overall efficiency and minimizes thermal losses. However, due to the pilot scale of this experimental setup, heat loss through the reactor walls is likely greater than in industrial installations. Factors contributing to this include the pilot reactor’s smaller scale than industrial plants, higher surface-to-volume ratio, and multiple sampling points, which collectively elevate thermal losses. Consequently, the cold gas and overall gasification efficiencies measured in these experiments may represent conservative values. In industrial reactors, which are generally larger with better insulation and lower surface-to-volume ratios, both cold gas and overall gasification efficiencies are likely to be higher due to minimized thermal losses.

The gasification technology utilized in this study is based on a fluidized bed system developed by EQTEC. The syngas characteristics,
<figure><img src="image_18.png" /><figcaption>Fig. 13. Energy balance and flow analysis on the gasifier unit for six selected test points.</figcaption></figure>

material flow, and energy analysis conducted on the semi-industrial
process demonstrate significant potential for upscaling the technology
to commercial applications. This potential is particularly evident with
the integration of advanced heat recovery systems for downstream ap-
plications, which can substantially enhance efficiency. The gasification
system's versatility makes it suitable for several commercial applica-
tions. The system is ideal for rural electrification or off-grid applications
due to its capacity to process locally sourced biomass feedstocks. This
makes it a viable solution for regions lacking centralized energy infra-
structure. Additionally, syngas generated can efficiently power Com-
bined Heat and Power (CHP) systems, which are well-suited for district
heating and industrial applications. These systems offer high overall
efficiency by utilizing both heat and power outputs. On a commercial
scale, the gasifier can be optimized by replacing air with oxygen/or
steam as the gasifying agent, enabling hydrogen-rich syngas production.
This would support the growing hydrogen economy, particularly in re-
gions transitioning to low-carbon energy systems. Furthermore, the
gasifier's ability to accommodate diverse biomass feedstocks, including
agricultural residues, forestry waste, energy crops, and municipal solid

waste, provides dual benefits of waste reduction and energy production. This feature aligns with circular economy principles and addresses waste management challenges.

Despite the promising applications, several barriers must be
addressed to successfully implement commercial-scale biomass gasifi-
cation systems. The substantial investment required for building and
commissioning industrial-scale gasifiers, including advanced gas clean-
ing and heat recovery systems, remains a significant hurdle [50].
Ensuring a steady, cost-effective supply of suitable biomass feedstock is
challenging, particularly in regions without established supply chains
for agricultural residues or forestry by-products. Several strategies can
be employed to overcome these barriers. Local governments can intro-
duce subsidies, tax breaks, and grants to incentivize investment in
biomass gasification technologies. Integration into national renewable
energy strategies can further promote commercialization of the gasifi-
cation system. Expanding end-use markets for syngas applications can
help to drive commercial uptake. Establishing a robust feedstock supply
chain by promoting the use of locally sourced agricultural and forestry
residues, specifically grown energy crops, and municipal waste can
mitigate supply chain challenges. Regular mixing of these feedstocks based on local availability and transportation logistics is also recommended [50–52]. The findings from this study underscore the significant potential of biomass gasification with a fluidized bed system to contribute to sustainable energy solutions. By addressing the outlined barriers through systematic adherence to proposed strategies, this technology can play a pivotal role in advancing renewable energy adoption and supporting the transition to a low-carbon economy.

## 4. Conclusion

This study provides a comprehensive evaluation of semi-industrial autothermal biomass gasification, bridging the gap between laboratory-scale research and full-scale industrial deployment. Through air gasification of woody biomass in a 200 kW bubbling fluidized bed gasifier, we assessed syngas composition, gasification performance metrics, energy efficiency, and system-wide energy distribution under varying equivalence ratios. The results demonstrated syngas composition varied significantly with operating conditions, with H₂ (6.4–11.3 %), CO (8.5–16.6 %), CH₄ (3.0–5.5 %), CO₂ (15.6–19.4 %), N₂ (48–62 %), C₂H₄ (0.8–1.7 %) and trace amounts of C₂H₂, C₂H₆, and C₆H₆, yielding a lower heating value of 3.9–7.1 MJ/Nm³. The system achieved cold gas efficiencies (CGE) of 51.8–71.4 % and carbon conversion efficiencies (CCE) of 78.6–98.5 %. Additionally, thermal energy recovery from the heat exchangers, bag filter, and cyclone separator contributed to an overall system efficiency of 79–93 %, highlighting the importance of integrating heat recovery mechanisms. The findings emphasize the critical role of thermal energy management in improving gasification systems sustainability and efficiency. The use of real-time heat recovery assessments (expressed as power output in kW) provides a dynamic perspective on energy flows within the system. This study underscores the feasibility of semi-industrial gasification as a scalable and efficient pathway for biomass-to-energy conversion, supporting advancements in renewable integration. Future work should explore further optimization and implementation of heat recovery strategies, integration with combined heat and power (CHP) systems, and catalytic tar cracking techniques to enhance syngas quality and operational efficiency in industrial applications.

## CRediT authorship contribution statement

P.O. Badu: Writing – original draft, Methodology, Data curation, Conceptualization. M. Debal: Writing – review & editing, Methodology, Data curation, Conceptualization. P. Girods: Writing – review & editing, Project administration, Methodology, Conceptualization. S. Aubert: Project administration, Methodology, Y. Rogaume: Project administration, Methodology, Funding acquisition, Conceptualization.

## Declaration of competing interest

The authors declare that they have no known competing financial interests or personal relationships that could have appeared to influence the work reported in this paper.

## Acknowledgments

The authors gratefully acknowledge the Lab of Excellence ARBRE for the support of the EA 4370 LERMAB (Carnot ICEEL member) grant overseen by the French National Research Agency as part of the "Investissements d'Avenir" program (ANR-11-LABX-0002-01, Lab of Excellence ARBRE) and EQTEC group for their collaboration and support in designing and installing the pilot plant for experimental work.

## Data availability

No data was used for the research described in the article.

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