Combustion and Power Generation Characteristics of a Thermoelectric Generator Fueled by Hydrogen-Enriched Compressed Natural Gas (HCNG)

Abstract

Hydrogen-enriched compressed natural gas (HCNG) is a promising transitional fuel for residential-scale distributed power, yet its impacts on direct-combustion thermoelectric generator (TEG) systems remain insufficiently quantified. In this study, a micro-scale TEG integrated with a commercially available self-aspirating household burner was experimentally investigated under thermal inputs of 700–2500 W and hydrogen blending ratios of 0–20 vol%, using open-loop water cooling to maximize heat rejection. The hot- and cold-side temperatures exhibited negligible variation with a hydrogen addition, and the maximum electrical output was essentially preserved across all blending ratios; at 2500 W the system delivered 75.8 W with a system efficiency of 3.03%. In contrast, hydrogen blending substantially reduced pollutant emissions: at 2500 W, CO decreased from 52.7 to 1 mg/m³ and CO₂ from 6.73% to 5.36% as the hydrogen fraction increased from 0 to 20 vol%. Meanwhile, combustion stability improved, indicated by a reduced coefficient of variation (0.77% → 0.49%). These results demonstrate that up to 20 vol% hydrogen blending can achieve significant emissions mitigation without compromising TEG power performance, supporting HCNG-fueled TEGs as a practical option for residential backup power.

1. Introduction

Combustion of fossil fuels releases large quantities of greenhouse gases. The power industry is a major contributor, accounting for approximately 40% of the world’s total greenhouse gas emissions [1]. Therefore, there is an urgent need for alternative solutions to reduce carbon emissions in the power sector. One solution is to use surplus renewable electricity for water electrolysis to produce green hydrogen. Mixing a certain proportion of hydrogen with natural gas (NG) yields a new type of gaseous fuel known as hydrogen compressed natural gas (HCNG). Due to the extensive infrastructure already in place for natural gas (NG), mixing hydrogen into existing pipelines enables efficient transportation and storage of hydrogen, offering a promising pathway for green hydrogen utilization. Urban gas systems (mainly referring to NG used for household, commercial, and district heating applications) account for approximately 40% of total NG demand. As such, they have a substantial capacity to accommodate HCNG. Many studies, mostly in the combustion engine field, have focused on the combustion of HCNG, exploring its combustion characteristics. For example, Shahid et al. [2] and De Simio et al. [3] investigated HCNG utilization in engine experiments, demonstrating that adding a certain amount of hydrogen reduces CO and unburned hydrocarbon emissions. On the other hand, through simulation research, Navarro et al. [4] and Zareei et al. [5] reached a similar conclusion: adding a certain amount of H₂ can reduce CO and CO₂ emissions. The aforementioned studies primarily focus on applying HCNG on engine applications. More than that, HCNG can also be used for generating power and heat via thermoelectric generators (TEGs), especially in extreme weather and occasionally grid failure conditions.

TEG is a type of distributed power source, operating based on the Seebeck effect [6]. TEG is characterized by simple structure, minimal moving parts, low maintenance requirements, and quiet operation. Micro- and mesoscale TEGs driven by gas combustion have attracted significant research interest. For example, He et al. [7] developed a disc-type combustor for hydrogen catalytic combustion in a micro-TEG, achieving an impressive power output of 1.87 W. Li et al. [8] designed a small-scale hydrogen catalytic combustion TEG that produced 20.7 W of power with an efficiency of 3.44%.

A literature survey shows that most research on TEGs employing hydrogen (enriched) fuel adopts catalytic combustion and direct combustion studies remain very rare. Due to hydrogen’s high flame speed and inherent instability, direct hydrogen combustion can easily cause flashback [8]. These challenges can be mitigated by blending hydrogen with methane or propane. Xie et al. [9] conducted premixed combustion of methane, hydrogen, and air in a micro porous media combustor, achieving an electrical output of 2.5 W and a system efficiency of 2.1%. Similarly, Ding et al. [10] investigated TEG using methane/hydrogen blended fuels in a micro porous media combustor, reporting an output of 57.8 W and an efficiency of 2.6% at a hydrogen mixing ratio of 30%. Although HCNG has been extensively studied in applications such as gas stoves [11], swirl burners [12], porous media burners [13], and internal combustion engines [14], the direct combustion-driven TEG using HCNG remains largely unexplored. Beyond hydrogen enrichment, other low-carbon gaseous additives, such as ammonia, have also been explored in microscale combustion-driven power systems. For example, Zhao et al. [15] investigated an ammonia/methane-fueled micro-thermal photovoltaic system and analyzed the electrical power output, energy efficiency, and NO/CO emissions. However, compared with ammonia-based fuels, hydrogen-enriched natural gas is more readily compatible with existing natural-gas infrastructure and household burner systems, which motivates the present investigation of HCNG-fueled thermoelectric generators.

In this study, we design and test a small-scale TEG powered by direct combustion of HCNG, aiming to evaluate its temperature profile, power output, efficiencies and emissions relative to NG. A small-scale, naturally aspirated gas stove burner was used to simulate the combustion behavior of HCNG under typical household conditions and to evaluate its effects on TEG performance. Currently, this study is the first to report on a HCNG-fueled TEG for household applications, innovatively using a household self-aspirating burner to combust hydrogen-enriched natural gas as the heat source for the TEG. It addresses the research gap of using direct HCNG combustion powered TEG systems in household applications.

2. Materials and Methods

2.1. The Setup of the TEG

The setup of the TEG is shown in Figure 1. The TEG consists of a frame, a gas mixer (GM), a combustion chamber, three self-aspirating burners, a pulse igniter, a heat collector, two water-cooled heat exchangers (HEXs), 12 thermoelectric modules (TEMs), and two clamps. Excluding the frame, the overall dimensions of the TEG are 65 × 100 × 260 mm, and it weighs 3.5 kg. The frame is made of stainless steel and is designed with a flue gas channel 150 mm in height to prevent air interference during flue gas measurement. For ease of understanding the experimental results, Figure 1b also illustrates the temperature measurement locations on the hot side and the compensation length ($L_1$) [16]. During assembly, a pressure of 1 MPa is applied to the clamps securing the thermoelectric modules (TEMs) and the water-cooled heat exchangers. This clamping pressure is critical for minimizing contact thermal resistance, thereby optimizing the thermoelectric generator’s (TEG) performance.

From a more specific perspective, the TEG comprises two subsystems: the combustion subsystem and the power generation subsystem. These subsystems will be discussed in detail in Section 2.1.1 and Section 2.1.2, respectively.

2.1.1. Combustion Subsystem

The combustion subsystem primarily consists of a fuel–air mixer, a combustion chamber, and three self-aspirating burners.

The combustion chamber, fabricated from stainless steel, is illustrated in Figure 1b. Its overall dimensions are 127.5 × 95 × 87 mm. Stainless steel was selected for the combustion chamber owing to its high-temperature oxidation resistance and corrosion resistance; moreover, its comparatively low thermal conductivity reduces conductive losses through the chamber walls, allowing a greater fraction of heat to be absorbed by the heat collector. Three self-aspirating burners are installed on the base. An electronic ignition needle is positioned above these burners and secured to the right-side panel. An observation window is incorporated into the front panel to facilitate flame monitoring. A quartz glass plate (47 × 47 × 4 mm) is mounted in front of this window. Ceramic fiber wool is packed between the combustion chamber wall and the quartz glass plate to ensure thermal insulation.

The gas mixer is shown in Figure 1c, with a volume of 100 mL and constructed from stainless steel. The gas mixer features three interfaces: two are for the inputs of hydrogen and NG respectively, and one is for the HCNG output. The interfaces are connected to polytetrafluoroethylene tubing via quick-connect fittings made of 316 stainless steel. Within the gas mixer, six baffles are configured with a center-to-center spacing of 25 mm between neighboring baffles. The hole positions on adjacent baffles differ slightly. The holes on the baffles have a diameter of 3 mm. The primary function of the baffles is to disrupt the laminar flow of the gases, thereby promoting the thorough mixing of NG and hydrogen.

As shown in Figure 1d, the self-aspirating burners utilized in this study are constructed from copper, with three burners located at the lowest section of the combustion chamber and interconnected by flanges. The gas nozzle possesses a diameter of 0.06 mm, through which gas is expelled and directed onto a buffer plate with a diameter of 9 mm, situated 15 mm above the nozzle. Furthermore, to facilitate optimal combustion, six principal air intake apertures, each measuring 3 mm in diameter, are included at the base of each burner. Combustion transpires above the buffer plate, as the gas flow ascends after impacting the buffer plate, hence enhancing the combustion process. The selection of this burner is justified by its widespread commercial availability and predominant use in household applications, making it an ideal candidate for investigating the applicability of HCNG-fueled TEGs in domestic settings.

2.1.2. Power Generation Subsystem

The power generation subsystem primarily consists of 12 thermoelectric modules (TEMs), a heat collector assembly, and two water-cooled heat exchangers.

The TEM used in this study is model TEG1-12708, manufactured by SAGREON Co., Ltd., Wuhan, China. It has dimensions of 40 × 40 × 3.8 mm. The physical properties, such as electrical resistivity and Seebeck coefficient, as well as the structural dimensions of the ceramic plates and legs, can be found in [16]. The TEM’s data sheet states that it can produce 10.8 W of electrical power with a thermoelectric conversion efficiency of 5.7% at hot and cold side temperatures of 523 K and 303 K, respectively.

The heat collector is manufactured using a CNC machine and is made of 6061-T4 aluminum alloy. It consists of two interleaved pin fin-assisted heat collection components and pin fins increase the effective contact area with the flue gas [17]. This improves heat collection and makes the temperature more uniform, which in turn helps TEMs better use their power-generation potential [18]. The dimensions of the heat collector are 128 × 95 × 168 mm. Each heat collector assembly has 139 interleaved pin fins that are 30 mm in length and 4 mm in diameter. The pin fins, fluid channels, and temperature uniformity plates in each heat collector are all manufactured from a piece of aluminum, which eliminates the negative impact on heat conduction that may arise from separately processing each structural component. Additionally, during the machining of the heat collector, six temperature measurement holes were integrated into the heat ends, facilitating accurate temperature measurement at the hot ends using surface-mounted thermocouples.

The water-cooled heat exchangers are made of 6063-T4 aluminum alloy, with dimensions of 80 × 120 × 15 mm. They contain six parallel, equidistant circular channels with a diameter of 7 mm. A graphite foil with a thickness of 0.3 mm is applied between the TEM and the heat collector or water-cooled heat exchangers. The graphite foil has a thermal conductivity of 18.4 W/m·K. Reducing the heat resistance at the contact is the goal of employing graphite foil [19].

2.2. Experimental Setup

Figure 2 illustrates the experimental setup of the TEG. The experiment requires an NG cylinder, a hydrogen gas cylinder, two mass flow controllers (MFCs), a water flow meter, a gas mixer, an electronic load instrument, several thermocouples, and a data acquisition instrument (DAQ). HCNG is used as fuel. The flow rates are controlled by the two mass flow controllers (MFC, Alicat MC-6SLPM-D, Japan Star Techno Co., Ltd., Osaka, Japan) with an accuracy of ±0.6% F.S. After passing through the MFCs, the hydrogen and NG are thoroughly homogenized within the gas mixer. An electronic load instrument was connected to the TEG system, and the power output characteristics under different resistive loads were measured by adjusting the load resistance. The model of the electronic load instrument is Prodigit 3305F (Prodigit Electronics Co., Ltd., Taiwan, China), with an accuracy of ±0.1% F.S. The medium entering the water-cooled heat exchanger is city tap water. A water flow meter regulates the flow rate of the cooling water (IkunK24, AQUA System Co., Ltd., Hikone, Japan) with an accuracy of ±4.0%. An open-loop cooling system is employed in this study. The objective of employing an open-loop cooling system in this experiment is to maximize electrical power output.

Temperatures are measured using K-type thermocouples, including the hot-end temperatures ($T_{\mathrm{h}1}$, $T_{\mathrm{h}2}$, $T_{\mathrm{h}3}$), cold-end temperatures ($T_{\mathrm{c}1}$, $T_{\mathrm{c}2}$), exhaust gas temperature ($T_{\mathrm{out}}$), ambient temperature ($T_{\mathrm{atm}}$), combustion chamber sidewall temperature ($T_b$), inlet water temperature ($T_{\mathrm{wt}1}$), and outlet water temperature ($T_{\mathrm{wt}2}$), with an accuracy of ±0.5%. An S-type thermocouple is used to measure the flame temperature ($T_f$), with a measurement accuracy of ±1.0%. All temperature signals were recorded using a data acquisition instrument (HY005, KODAMA GLASS, Minamikameicho, Osaka, Japan). A flue gas analyzer (Testo340, Senze Instruments Pte Ltd., 28F Penjuru Cl, Singapore) is used to measure the concentrations of CO, ${\mathrm{CO}}_2$, and ${\mathrm{NO}}_{\mathrm{x}}$ in exhaust gases. The detailed specifications of the above-mentioned equipment are listed in

Table 1. Detail specifications for the equipment used in the experiment.

Equipment	Model	Accuracy	Unit	Measurement
Mass Flow Controller	Alicat MC-20SLPM-D	±0.6%	LPM	$V_{\mathrm{fuel}}$
Electronic Load Instrument	Prodigit 3305F	±0.1%	W	$U$, $I$, $P_{\mathrm{tot}}$
Water Flow Meter	IkunK24	±4.0%	L/min	$V_w$
Type-K Thermocouple, TC Ltd., Uxbridge, UK	Type-K	±0.5%	K	$T_h$,$T_c$,$T_{\mathrm{out}}$,$T_{\mathrm{atm}}$,$T_b$,$T_{\mathrm{wt}}$
Type-S Thermocouple, TC Ltd., Uxbridge, UK	Type-S	±1.0%	K	$T_f$
Data Acquisition Instrument	HY005	\	\	\
Flue Gas Analyzer	Testo 340	±2.0, ±2.0, ±0.3	mg/m3, mg/m3, %	CO, ${\mathrm{NO}}_{\mathrm{x}}$, ${\mathrm{CO}}_2$

.

2.3. Parameter Definitions

The fuel is a mixture of hydrogen and NG. The hydrogen blending ratio (${\epsilon }_{\mathrm{hydrogen}}$) in the fuel can be calculated using the following equation:

$${\epsilon }_{\mathrm{hydrogen}}={\displaystyle \frac{V_{H_2}}{V_{\mathrm{C}{\mathrm{H}}_4}+V_{H_2}}}\times 100\%,$$

(1)

where $V_{\mathrm{C}{\mathrm{H}}_4}$ and $V_{H_2}$ represent the volumetric flow rates of NG and hydrogen, respectively. The total chemical energy of the mixture corresponds to the input power ($P_{\mathrm{in}}$) of the TEG, denoted as:

$$P_{\mathrm{in}}=q_{\mathrm{C}{\mathrm{H}}_4}+q_{H_2}=V_{\mathrm{C}{\mathrm{H}}_4}\times {\mathrm{L}\mathrm{H}\mathrm{V}}_{\mathrm{C}{\mathrm{H}}_4}+V_{H_2}\times {\mathrm{L}\mathrm{H}\mathrm{V}}_{H_2},$$

(2)

where ${\mathrm{L}\mathrm{H}\mathrm{V}}_{\mathrm{C}{\mathrm{H}}_4}$ and ${\mathrm{L}\mathrm{H}\mathrm{V}}_{H_2}$ are the lower heating values of the NG and hydrogen, respectively. The ${\mathrm{L}\mathrm{H}\mathrm{V}}_{\mathrm{C}{\mathrm{H}}_4}$ and ${\mathrm{L}\mathrm{H}\mathrm{V}}_{H_2}$ are 32.6 MJ/m³ and 9.9 MJ/m³, respectively [20]. A portion of the fuel’s energy is not fully released due to partial combustion. The heat released during combustion is partly collected by the heat collector, while a portion is lost via natural convection and thermal radiation ($q_{\mathrm{loss}1}$), which is unavoidable. The other part involves the heat flux in the flue gas and ($q_{\mathrm{flue}}$) in the exhaust gas. Therefore, the heat flux density collected ($q_{\mathrm{HC}}$) [21] by the heat collector can be derived as:

$$q_{\mathrm{HC}}=q_{\mathrm{comb}}-q_{\mathrm{flue}}-q_{\mathrm{loss}1}=P_{\mathrm{in}}\times {\eta }_{\mathrm{comb}}-q_{\mathrm{flue}}-q_{\mathrm{loss}1},$$

(3)

where $q_{\mathrm{comb}}$ represents the energy released by the fuel combustion and ${\eta }_{\mathrm{comb}}$ represents the combustion efficiency. The heat collected by the heat collector is partially converted into electricity ($P_{\mathrm{tot}}$) by the TEM based on the Seebeck effect. Another portion is lost through convective heat transfer and thermal radiation ($q_{\mathrm{loss}2}$). The largest portion is transferred via the cooling water of the water-cooled heat exchanger ($q_w$). Therefore, $q_{\mathrm{HC}}$ can also be defined as:

$$q_{\mathrm{HC}}=P_{\mathrm{tot}}+q_{\mathrm{loss}2}+q_w=P_{\mathrm{tot}}+q_{\mathrm{loss}2}+c_w\times {\rho }_w\times V_w\times \left(T_{\mathrm{wt}2}-T_{\mathrm{wt}1}\right),$$

(4)

where $c_w$, ${\rho }_w$, $V_w$, $T_{\mathrm{wt}1}$, and $T_{\mathrm{wt}2}$ are the specific heat capacity, density, volume flow rate of the tap water, the inlet temperature, and the outlet temperature, respectively.

The natural convection heat transfer and thermal radiation from the combustion chamber and the heat collector ($q_{\mathrm{loss}1}$ and $q_{\mathrm{loss}2}$) are inevitable. $q_{\mathrm{loss}1}$ and $q_{\mathrm{loss}2}$ can be calculated as follows:

$$q_{\mathrm{loss}1}=q_{\mathrm{conv}}+q_{\mathrm{rad}}=h_{\mathrm{air}}\times A_C\times \left(T_{\mathrm{b},\mathrm{av}}-T_{\mathrm{atm}}\right)+\epsilon \times \sigma \times A_C\times \left(T_{\mathrm{b},\mathrm{av}}^4-T_{\mathrm{atm}}^4\right),$$

(5)

$$q_{\mathrm{loss}2}=q_{\mathrm{conv}}+q_{\mathrm{rad}}=h_{\mathrm{air}}\times A_{\mathrm{HC}}\times \left(T_{\mathrm{h},\mathrm{av}}-T_{\mathrm{atm}}\right)+\epsilon \times \sigma \times A_{\mathrm{HC}}\times \left(T_{\mathrm{h},\mathrm{av}}^4-T_{\mathrm{atm}}^4\right),$$

(6)

where $h_{\mathrm{air}}$, $\epsilon$, $\sigma$, $A_C$, $A_{\mathrm{HC}}$, $T_{\mathrm{b},\mathrm{av}}$, $T_{\mathrm{h},\mathrm{av}}$, and $T_{\mathrm{atm}}$ are the natural convection heat transfer coefficient, emissivity, Stefan–Boltzmann constant, combustion chamber sidewall area, collector sidewall area, average temperature of the combustion chamber sidewall, average hot-side temperature, and ambient temperature, respectively.

The thermal loss of exhaust gas ($q_{ext}$) is obtained as follows [16]:

$$q_{ext}=c_p\times {\rho }_{ext}\times U_{ext}\times A_{ext}\times \left(T_{ext}-T_{atm}\right)$$

(7)

where $c_p$, ${\rho }_{ext}$, $U_{ext}$, and $A_{ext}$ are the exhaust gas heat capacity, density, velocity, and area of the exhaust gas outlet, respectively. $T_{ext}$ and $T_{atm}$ are the exhaust gas temperature from the TEG and the atmosphere temperature, respectively.

Based on Equations (1)–(5), several key efficiencies can be derived. First, the overall efficiency (${\eta }_{\mathrm{sys}}$) is defined as the ratio of the electrical energy generated by the TEM to the input power, as follows:

$${\eta }_{\mathrm{sys}}={\displaystyle \frac{P_{\mathrm{tot}}}{P_{\mathrm{in}}}}$$

(8)

Combining Equations (1) and (2), the heat collection efficiency (${\eta }_{\mathrm{heat}}$) is defined as the ratio of the heat collected by the heat collector to the heat generated by the combustion of the fuel, and can be calculated as:

$${\eta }_{\mathrm{heat}}={\displaystyle \frac{q_{\mathrm{HC}}}{P_{\mathrm{in}}{\eta }_{\mathrm{comb}}}}$$

(9)

Another important parameter, the thermoelectric efficiency (${\eta }_{\mathrm{TE}}$), is defined as the ratio of the electrical energy generated by the TEM to the heat collected by the heat collector, and is given by:

$${\eta }_{\mathrm{TE}}={\displaystyle \frac{P_{\mathrm{tot}}}{q_{\mathrm{HC}}}}$$

(10)

To determine the changes in the flame after hydrogen blending, the average flame temperature, standard deviation, and coefficient of variation can be defined as [21]:

$$T_{flame,ave}={\displaystyle \frac{\sum T_{flame}}{N}}$$

(11)

$$\mathrm{SD}=\sqrt{{\displaystyle \frac{{\sum \left(T_{flame}-T_{flame,ave}\right)}^2}{N}}}$$

(12)

$$\mathrm{COV}={\displaystyle \frac{\mathrm{SD}}{T_{flame,ave}}}$$

(13)

where $N$ is the total number of samples taken. The flame temperature is the result of measuring the temperature at the centerline of the flame multiple times and taking the average value.

2.4. Experimental Case and Procedure

Table 2. Experimental cases for different input power levels at hydrogen blending ratios of 0% and 5%, and for various hydrogen blending ratios at an input power of 2500 W.

–	${\mathit{V}}_{\mathbf{C}{\mathbf{H}}_4}$ (L/min)	${\mathit{V}}_{{\mathit{H}}_2}$ (L/min)	${\mathit{\epsilon }}_{{\mathit{H}}_2}$ (%)	${\mathit{P}}_{\mathbf{in}}$ (W)	${\mathit{R}}_{\mathbf{ld}}$ (Ω)	${\mathit{V}}_{\mathit{w}}$ (L/min)	Cooling Mode
1	0.92–4.60	0	0%	700–2500 (600 *)	3–49	4.3	Open loop
2	0.91–4.53	0.05–0.24	5%	500–2500 (200 *)	3–49	4.3	Open loop
3	4.56	0.14	3%	2500	3–49	4.3	Open loop
4	4.48	0.39	8%	2500	3–49	4.3	Open loop
5	4.43	0.55	11%	2500	3–49	4.3	Open loop
6	4.38	0.71	14%	2500	3–49	4.3	Open loop
7	4.33	0.89	17%	2500	3–49	4.3	Open loop
8	4.28	1.07	20%	2500	3–49	4.3	Open loop

* Represents the step size.

presents the experimental test conditions. Two sets of experiments were conducted. The hydrogen blending ratios examined in this study were selected with reference to the maximum allowable limits set by different countries. As shown in Figure 3, comparing the maximum permitted ratios in Italy [22], France [23], the United Kingdom [24], and China [25] led us to cap the blending level at 20%, thereby enhancing the generalizability of the study. Furthermore, a study on gas interchangeability in reference [26] suggests an upper blending limit of 23%. Experimental investigations on HCNG combustion using residential gas stoves [27] show that when the hydrogen blending ratio is below 20%, both thermal efficiency and flame stability meet regulatory requirements. Taken together, these studies suggest that most household gas appliances can operate safely with HCNG containing up to 20%.

To ensure repeatability and accuracy, each experimental group was repeated three times. The experimental sets are as follows: (1) Examining the effect of hydrogen blending ratios (0% and 5%) on TEG performance across input powers ranging from 500 W to 2500 W. (2) Examining the effect of various hydrogen blending ratios (0% to 20%) on TEG performance at a fixed input power of 2500 W. The 5% hydrogen blending ratio was selected as the hydrogen adoption baseline, due to the findings of Chen et al. [13]. They found that at a 5% hydrogen blending ratio, the maximum deviation between actual and rated heat loads was 9.3%, which meets the Chinese national standard requirement (within ±10%). Additionally, no combustion instabilities such as flashback or flameout were observed.

The experimental procedure involves several steps:

• Verify ambient temperature and humidity; open the cooling water valve to ensure the cold-side temperature maintains a constant value.
• Set the flow rates of NG and hydrogen; open the fuel supply valve. Simultaneously press and hold the ignition electrode button until combustion stabilizes.
• Wait until the hot-side temperature stabilizes at a constant value, then record relevant data.
• After completing measurements, first close the fuel supply valve and allow system cooling. Then close the cooling water valve to conclude the experiment or repeat the procedure starting from step (2).

Variations in ambient temperature and humidity can influence both the combustion characteristics of the fuel and the heat dissipation at the cold side of the TEG. To minimize experimental errors, ambient temperature and humidity were maintained at 25–26 °C and 50–60%, respectively, for each test. Additionally, all experiments were conducted in strict accordance with standardized procedures, and data were recorded only after the TEG reached a steady-state condition.

2.5. Error Analysis

Table 3. Errors of the measured and derived parameters.

Parameter	Error	Parameter	Error
$U$ (voltage)	±0.1%	$T$ (type K thermocouple)	±0.5%
$I$ (current)	±0.1%	$T$ (type S thermocouple)	±1.0%
$V_w$ (tap water flow rate)	±4.0%	$V_{\mathrm{fuel}}$ (volumetric flow rate)	±0.6%
$P$ (electrical power)	±0.14%	${\eta }_{\mathrm{TE}}$	±3.11%
${\eta }_{\mathrm{sys}}$	±0.17%	${\eta }_{\mathrm{heat}}$	±3.92%

presents the standard errors of the measured parameters, calculated based on the measurement precision of the instruments listed in Table 1 and the standard error propagation function [28]. The associated uncertainties for electrical power, system efficiency, thermal collection efficiency, and thermoelectric efficiency are 0.14%, 0.17%, 3.92%, and 3.11%, respectively.

3. Results and Discussions

3.1. Flame Stability

To investigate the changes in the burner flame after hydrogen blending, the coefficient of variation (COV) was introduced to assess the variation in flame stability. The input power was kept constant at 2500 W, and 10 measurements were taken along the central axis of the self-aspirating burner to obtain the average flame temperature. The standard deviation (SD) and COV changes were then calculated using Formulas (12) and (13).

As shown in Figure 4, with the input power fixed at 2500 W, as the hydrogen blending ratio increased from 0% to 20%, the average molecular weight of the mixed fuel decreased, which led to an increase in the fuel injection speed. This caused a reduction in the fuel residence time, resulting in a downward trend in the average flame temperature along the axis, specifically decreasing from 1305 K to 1193 K, a drop of approximately 8.6%. Correspondingly, the SD value also decreased with the increase in hydrogen blending ratio, dropping from 10.07 to 5.82, a reduction of about 42.2%. The COV value also continuously decreased from 0.77% to 0.49%, demonstrating that as the hydrogen blending ratio increased, the flame stability improved.

3.2. Temperature Distribution

Figure 5a,b illustrates the temperature distributions at hydrogen blending ratios of 0% and 5% and across input powers ranging from 700 W to 2500 W. Firstly, it is evident that the cold-side temperature exhibits minimal variation. Specifically, T_c1 changes from 291 K to 296.3 K at a 0% hydrogen blending ratio and from 291 K to 298 K at a 5% ratio. Similarly, T_c2 increases from 292 K to 299 K (0% H₂) and from 293 K to 301 K (5% H₂). The slight elevation of T_c2 compared to T_c1 occurs because the heat exchanger (HEX) absorbs heat from the heat collector and transfers it to the cooling water. Consequently, the HEX outlet temperature (T_c2) exceeds the inlet temperature (T_c1). However, due to the high heat capacity of water and the open-loop cooling configuration, the resulting temperature increases at both T_c1 and T_c2 remain limited.

Secondly, the hot-side temperatures (T_h1, T_h2, T_h3) increase with rising input power. This occurs because higher input power releases more heat than can be absorbed by the heat collector. Furthermore, a decreasing trend is observed from T_h1 to T_h3. This temperature gradient arises due to thermal convection and radiation within the combustion chamber. These heat transfer mechanisms result in less heat being absorbed by the upper regions of the heat collector compared to the lower regions, consequently causing the observed temperature to decrease along its vertical height.

Figure 5c presents average temperatures of the hot-side, cold-side, and flame at a fixed input power of 2500 W and hydrogen blending ratios ranging from 0% to 20%. It is observed that with increasing hydrogen blending ratios, both hot-side (T_h) and cold-side (T_c) temperatures exhibit minimal variation. However, the peak flame temperature increases slightly with higher hydrogen blending ratios. This phenomenon is attributed to hydrogen’s higher combustion velocity and elevated adiabatic flame temperature compared to natural gas (NG). Consequently, hydrogen enrichment leads to an increase in the peak flame temperature.

Consequently, the experimental results demonstrate that the addition of hydrogen does not reduce thermoelectric module (TEM) operating temperatures. Conversely, flame temperature exhibits a measurable increase. These findings provide preliminary evidence for HCNG compatibility with household thermoelectric generators (TEGs). Power generation characteristics will be analyzed in Section 3.3.

3.3. Pollutant Emissions

Pollutant emissions pose critical environmental and health challenges in both industrial and residential applications. Compounds such as carbon monoxide (CO), carbon dioxide (CO₂), and nitrogen oxides (NO_x) contribute significantly to atmospheric degradation and climate change through greenhouse gas effects. In residential settings, these emissions present direct health hazards—particularly CO poisoning risks. Therefore, comprehensive emission analysis is essential for evaluating household thermoelectric generator (TEG) safety.

Figure 6a,b illustrates the pollutant emissions at hydrogen blending ratios of 0% and 5% and across input powers ranging from 700 W to 2500 W. Experimental results indicate significantly increased pollutant emissions with rising input power. Specifically, CO₂ emissions rise due to greater NG combustion at higher power levels, where methane oxidation (CH₄ + 2O₂ → CO₂ + 2H₂O) generates carbon dioxide. CO concentrations, being intermediate combustion products, serve as indicators of reaction completeness; increased fuel flow promotes incomplete combustion, resulting in elevated CO emissions. NOₓ formation exhibits strong temperature dependence as higher power releases more thermal energy, consequently increasing NOₓ production through thermal mechanisms. More significantly, experimental data demonstrate that a 5% hydrogen blending ratio exhibits lower CO and CO₂ emissions compared to NG. This indicates that the addition of hydrogen can significantly reduce these emissions.

Figure 6c illustrates pollutant emissions at a fixed input power of 2500 W with hydrogen blending ratios ranging from 0% to 20%. The data demonstrate an inverse relationship between hydrogen blending ratio and emission levels: increased hydrogen blending corresponds to decreased CO and CO₂ outputs [29]. For example, at 0% hydrogen blending, CO emissions measure 52.7 mg/m³ and CO₂ concentration is 6.73%. At 20% hydrogen blending, these values decrease to 1 mg/m³ and 5.36% respectively. This phenomenon can be attributed to the fact that hydrogen is a zero-carbon fuel, and its combustion primarily produces water vapor (H₂O). In a mixed combustion process, the addition of hydrogen promotes more complete combustion; thus reducing the formation of carbon monoxide (CO). CO is an intermediate product of hydrocarbon combustion (such as natural gas) when combustion is incomplete, typically occurring in oxygen-deficient conditions. As hydrogen is introduced, its high combustion speed and lower ignition energy help improve combustion efficiency, thereby reducing CO formation. For the same reason, since hydrogen itself contains no carbon, it also reduces carbon dioxide emissions, which aligns with the observed results.

As for NO_x, its emissions generally increase with the hydrogen blending ratio [30]. However, in this experiment, NO_x emissions fluctuated and did not show a clear increasing trend. Notably, NO_x emissions did not exhibit the expected surge even at 20% hydrogen blending. This phenomenon is attributed to the synergistic effect of two mechanisms inherent to the system. First, the intensive thermal quenching by the pin-fin heat collector acts as a dominant thermal sink. Rapid heat extraction suppresses the bulk flame temperature below the critical threshold for thermal NO_x formation [31]. Secondly, a momentum-induced leaning occurs within the self-aspirating burner. The high-velocity hydrogen jet enhances air entrainment, shifting the local combustion towards a leaner condition [22]. These combined effects effectively counteract the adiabatic temperature rise of hydrogen, preventing a net increase in emissions. It can be concluded that this TEG device, with a hydrogen blending ratio of up to 20%, not only reduces CO and CO₂ emissions but also keeps NO_x emissions within an acceptable range, which is crucial for the health of households using this TEG in the future.

3.4. Power Generation and Efficiency

Figure 7a,b illustrates power output variation ranging from 700 W to 2500 W at hydrogen blending ratios of 0% and 5%, respectively. The data demonstrates a proportional increase in output power with increasing input power. Specifically, at 0% hydrogen blending and 700 W input, output power measures 9 W; at 2500 W input, output increases to 75.8 W. This trend persists at 5% hydrogen blending (Figure 7b), where output reaches 8.9 W (700 W input) and 75.8 W (2500 W input).

From Figure 7a,b, it is evident that hydrogen blending ratios do not significantly influence output power. This represents a critical factor for thermoelectric generators (TEGs) utilizing HCNG fuel. Figure 7c illustrates output power, thermoelectric (TE) efficiency, and system efficiency at a fixed input power of 2500 W with hydrogen blending ratios ranging from 0% to 20%. It is observed that increasing hydrogen blending ratios produce no significant change in efficiency or output power, with values fluctuating near 75 W. This phenomenon aligns with expectations, since hydrogen blending does not offer extra heating values or substantially alter hot-side or cold-side temperatures (as shown in Section 3.1), consequently maintaining stable output power. This represents a significant finding, as the addition of hydrogen preserves the power output capability of thermoelectric generators (TEGs). Consequently, HCNG proves superior to NG for household TEG applications. While maintaining equivalent power generation, HCNG emits far less CO/CO₂—benefiting both indoor air quality and environmental sustainability. During extreme weather events or grid failures, TEGs can provide continuous power supply. Furthermore, with hydrogen’s potentially lower cost in certain markets, operational expenses may be reduced.

3.5. Energy Flow Analysis

The application of HCNG in the residential sector is currently focused primarily on appliances such as gas stoves and water heaters. The evaluation criteria for these applications mainly include thermal load, thermal efficiency, combustion temperature, combustion stability, and pollutant emissions [32]. This study differs from traditional research based on gas stoves. It evaluates the impact of HCNG on the performance of TEG, exploring and expanding the application range of HCNG in energy conversion. This not only establishes a theoretical framework and experimental support for its utilization in the civilian domain but also serves as a basis for the future formulation of pertinent standards. After hydrogen is doped into NG, both the combustion temperature and combustion rate are significantly enhanced. This is because hydrogen has a lower ignition energy and faster flame propagation speed compared to NG, which greatly increases the combustion rate of the HCNG mixture. The fast combustion rate of the HCNG mixture leads to more rapid and concentrated energy release, directly resulting in higher combustion temperatures.

Figure 8 shows an energy flow Sankey diagram, indicating the energy distribution and potential directions for further optimization when the input power is 2500 W and the hydrogen blending ratio is 5%. The TEG system demonstrates an aggregate energy loss of 718.3 W, comprising incomplete fuel combustion and exhaust gas dissipation. Current instrumentation limitations prevent precise quantification of individual contributions. Nevertheless, we can perform a preliminary calculation using the relevant formulas. When the input power is 2500 W and the output power is 74.8 W (2.99%), the energy carried away by the cooling water (water-cooled exchanger) can be calculated using Formula (4), resulting in 1527 W (61.08%). Then, using Formulas (5) and (6), the losses from natural convection heat transfer and thermal radiation from the combustion chamber and heat collector (convection and radiation) can be determined, resulting in 179.9 W (7.2%). Finally, by substituting the relevant parameters obtained from the experiment into Formula (7), the energy lost through exhaust gas heat (fuel gas heat loss) can be calculated, resulting in 593.3 W (25.04%). This allows us to determine the energy attributed to incomplete combustion, which is approximately 125 W. As a result, the energy distribution during the TEG operation is initially determined, with incomplete combustion accounting for about 125 W of the total loss, while the remaining 593.3 W is attributed to exhaust gas thermal dissipation. Natural convective heat transfer and thermal radiation account for approximately 7.2% of the energy loss, which is unavoidable. However, improvements in combustor design, combustion strategies, and heat collector optimization could enhance both combustion and heat collection efficiency. The water-cooled heat exchanger transfers approximately 61% of the total energy, while only 2.99% is converted into electrical power. This indicates that a significant portion of the energy remains unutilized. Therefore, maximizing the recovery of energy lost through exhaust gases and the water-cooled heat exchanger is crucial for improving TEG performance. Although NG doped with a low proportion of hydrogen demonstrates good performance in TEG, particularly in reducing carbon emissions, this is mostly due to the decrease of carbon content in the fuel after hydrogen doping. Specifically, hydrogen is a zero-carbon fuel, and its combustion products are mainly water vapor. The addition of hydrogen decreases the carbon content of the fuel, hence cutting carbon dioxide emissions. This is of significant importance for achieving cleaner energy use and reducing greenhouse gas emissions. Nonetheless, regarding thermal efficiency, the system efficiency of the TEG remains suboptimal. This is mostly attributable to the fact that most commercially available thermoelectric materials have a thermoelectric efficiency of less than 6% [33]. Therefore, despite the advantages of the TEG in terms of carbon emission reduction and energy conversion, its relatively low system efficiency remains an important challenge that needs to be addressed.

3.6. Quantitative Comparison with Prior Studies

The primary objective of this study was to design a TEG suitable for household emergency use and, in view of the anticipated wider adoption of HCNG, we choose it as the fuel. We first examined the temperature and emissions characteristics of a commercially available, self-aspirating burner firing HCNG, and found that hydrogen addition increased the flame temperature without altering the TEG’s hot- and cold-side temperatures. More importantly, relative to NG, HCNG combustion produced markedly lower pollutant emissions, including CO and CO₂. As shown in Figure 9a, our CO emissions are compared with related experiments; the reduction is attributed to more complete combustion and to hydrogen partially displacing NG in oxidation, thereby suppressing CO formation. Given that household environments are often relatively confined, such minimal emissions are critical for resident health and life safety. Meanwhile, considering the continued rise in residential electricity demand, we compared output power across reported TEG studies and found that the present system delivers comparatively high output, enabling it to provide as much electrical support as possible to a household during emergencies.

4. Conclusions

This study designs and tests a TEG driven by the direct combustion of HCNG, filling the research gap regarding the use of HCNG as a fuel for driving TEG. The TEG system is characterized by high output power, low carbon emissions, and stable operation. Following comprehensive study and discussions on the experimental outcomes, the subsequent conclusions are derived:

• At a fixed input power, progressively increasing the hydrogen blending ratio from 0% to 20% produced no discernible change in the heat collector hot- and cold-side temperatures, although the peak flame temperature increased. This finding indicates that the employment of HCNG in domestic gas burners does not compromise the system’s heat-transfer performance.
• Hydrogen addition can effectively reduce CO and CO₂ pollutant emissions. For example, at an input power of 2500 W, CO and CO₂ emissions at a 0% hydrogen blending ratio were 52.7 mg/m³ and 6.73%, respectively; when the hydrogen blending ratio was increased to 20%, these values decreased to 1 mg/m³ and 5.36%, respectively.
• As the hydrogen blending ratio increases, the average flame temperature gradually decreases, from 1305 K to 1193 K, a drop of about 8.6%. At the same time, flame stability increases, with the coefficient of variation dropping from 0.77% to 0.49%. This indicates that as the hydrogen blending ratio increases, the stability of the flame improves progressively.
• The power output of TEG fueled by HCNG is comparable to that achieved with NG. At an identical input power of 2500 W, the output power at a 0% hydrogen blending ratio was 75.8 W, and at a 20% hydrogen blending ratio it likewise remained 75.8 W. These results show that it is completely feasible to use HCNG to drive TEG.