Inline induction heating for high-pressure fuel-based testing applications

Abstract

Induction heating offers numerous advantages over conventional methods for heating high-pressure fluids in hydraulic testing applications. It simplifies complex processes, mitigates the risks associated with high-pressure environments, and enhances overall safety. The design of an efficient induction heating system requires a clear understanding of its key components, including the resonant tank circuit, impedance matching and isolation stage, high-frequency inverter, driver circuitry, and controller. The configuration of the resonant circuit plays a crucial role in device ratings and control methods. To meet the demanding requirements of switching devices like IGBTs, the driver circuit must ensure high immunity to interferences and failure. Effective system design allows the heating circuit to operate close to resonant switching, minimizing switching losses and cooling requirements. This paper presents the design and development of a 4-kW induction heater specifically tailored for inline heating of aviation fuel under high pressure. The system aims to achieve a precise temperature profile for testing line-replaceable units (LRUs) used in aircraft. The design methodology encompasses the selection of the resonant tank circuit configuration, impedance matching and isolation stage, high-frequency inverter, driver circuitry, and control strategies. Through extensive simulations in LTspice and experimental measurements, power losses in different components of the power conversion circuits are assessed, and overall system efficiency is evaluated. The results demonstrate the proposed induction heating system’s remarkable energy efficiency, reduced switching losses, and improved reliability. Furthermore, a comprehensive comparison with existing designs and industry norms is presented, highlighting the distinct advantages of the proposed system in terms of cost-effectiveness, energy efficiency, material utilization, and processing time. This research contributes valuable insights into the design and optimization of induction heating systems for high-pressure fluid heating applications, providing a practical and efficient solution for the aviation industry.

Introduction

High-temperature testing is a crucial requirement for qualifying aircraft engine system hydraulic line-replaceable units (LRUs), including servo valves, actuators, filters, and more. These components often operate with aviation fuel as the working fluid and undergo testing with aviation fuel at high pressures, following specific temperature profiles as illustrated in Fig. 1. These temperature profiles aim to simulate the operational conditions and endurance of the aircraft over several flight hours.

Fig. 1

Typical fuel temperature profile for testing aircraft engine system hydraulic LRUs

Inline heating of the fuel is a preferred approach to ensure a continuous supply of dynamically heated fuel while minimizing energy expenses and the need for hot reservoirs. However, achieving the required temperature profiles through heating fuel at high pressures presents several challenges. Traditional immersion heating elements cannot withstand the high pressures inside the hydraulic lines, and their rupture within the system would pose significant safety hazards.

In contrast, induction heating emerges as a superior choice due to its non-invasive nature, electrical isolation, cleanliness, and several advantageous features such as flexibility, controllability, fast heating, repeatability, and safety [1,2,3,4,5]. This paper focuses on elaborating the design approach of a 4-kW induction heater specifically developed for inline heating of aviation fuel at a pressure of 84 bar and a flow rate of 1 lpm, incorporating closed-loop control.

The system’s basic block diagram, as depicted in Fig. 2, illustrates the interaction between the main process controller, the PID controller responsible for temperature regulation, and the induction heater, which is controlled using an ON/OFF strategy. The command signals to the PID controller are suitably adjusted to achieve the temperature profile shown in Fig. 1.

Fig. 2

Overall scheme of induction heating control with temperature feedback

Methods

Induction heating technique is widely used in numerous low- to high-power applications due to the various advantages it offers. The basic approach in the implementation of an induction heating system involves producing high-frequency alternating magnetic flux to generate heat in a conductive workpiece through the eddy currents induced in the workpiece. The technique utilizes the skin effect that forces the current to be concentrated near the surface of the object, thereby increasing the effective resistance of the workpiece producing more heat [6]. The thickness of this region where the current gets concentrated or the skin depth is given by [1, 2]

$$\updelta =\sqrt{\frac{2\uprho }{2\mathrm{\pi f\mu }}} (m)$$

(1)

where

ρ is electrical resistivity (Ω m)

f is the frequency (Hz)

μ is the magnetic permeability (H/m)

Heat pipe

In the case of a solid workpiece made of steel, the heat generated near the outer surface of the workpiece gets conducted to the inner portion of the workpiece to make it uniformly hot. However, during active heating through the induction technique, there exists a temperature gradient from the outer surface of the workpiece to its center due to the low specific heat capacity of the steel resulting in slow heat transfer. Instead, if the workpiece is a steel pipe with a thickness comparable with the skin depth, and carries a liquid with high specific heat capacity, such as aviation fuel, or even water, the heat produced near the surface of the workpiece can easily get conducted to the liquid inside, making it hot.

The heating pipe should however be sufficiently long to prevent excessive heating power focused on a small surface area of the heating pipe and causing rupture at high pressure due to heat.

The heating arrangement comprises a low carbon steel pipe with a 22-mm outer diameter and 2 mm thickness, which acts as the workpiece and is surrounded by a 19-turn helical copper coil with a pipe diameter of 8 mm and wall thickness of 2 mm. The length of the steel pipe under the coil is around 450 mm. As the copper coil is excited with a high-frequency alternating current, heat is induced through eddy currents near the outer surface of the pipe which is conducted across the thickness of the pipe and then transferred to the fuel that is passing through the pipe. Figure 3 shows the simulated view of this arrangement with a power input of 4 kW fed for a period of 5 s.

Fig. 3

Simulated view of the induction heated 22 mm dia steel pipe for inline fuel heating application with CENOS

Power circuit

The power circuit for the induction heater comprises of a high-frequency inverter as shown in Fig. 4 [7]. The DC voltage input to the inverter is obtained from a single-phase AC power supply through an AC voltage controller ${T}_{1}$ and a bridge rectifier. The inverter is operated at a constant frequency, while the AC voltage controller is used for implementing the closed-loop PID control.

Fig. 4

Power circuit of the induction heater with AC power supply, AC voltage controller, bridge rectifier, and high-frequency inverter

Capacitor ${C}_{1}$ is required to absorb the reactive currents from the output of the inverter without allowing high ripple voltages in the DC link [8]. However, as the induction heating inverter operates at high frequency and at resonance, with almost negligible reactive currents, a low DC link capacitance is usually sufficient.

The value of capacitance is calculated from the basic capacitor equation

$${i}_{{c}_{1}}={C}_{1}\frac{d{v}_{{c}_{1}}}{dt}$$

(2)

where ${i}_{{c}_{1}}$ is the change in current (A),

${C}_{1}$ is the capacitance (F),

$d{v}_{{c}_{1}}$ is the change in capacitor voltage (V), and

$dt$ is the change in time (s).

Rearranging (2),

$${C}_{1}=\frac{{i}_{{c}_{1}}}{2fd{v}_{{c}_{1}}}$$

(3)

where $f$ is the frequency (Hz).

Considering the peak ripple current as 10 percent of 15 A load current and 2 V as allowable ripple voltage at the DC link when operating at 37 kHz frequency, the DC link capacitance is given by

$${C}_{1}=\frac{1. 5}{2\times 37\times { 10}^{3}\times 2} =10. 1 \mu F$$

Polymer capacitor bank of 12 µFis used for the DC link due to its fast dynamic response and durability. IGBTs are ideal as switching devices for high-power resonant switching applications such as induction heating. They are capable of handling large power levels with low power gate drive [1, 7]. The IGBTs Q ₁, Q ₂, Q ₃, and Q ₄ are essentially fast-switching devices capable of operating till 100 kHz [7]. Their switching losses increase considerably beyond the frequency of 50 kHz with non-zero switching, which may prove to be highly inefficient. The IGBTs should be rated for at least 400 V, considering the peak voltage of 325 V DC link voltage at no load and possible overvoltages due to reactive currents from the load. Current rating of the IGBTs should be at least 5 to 7 times higher than the maximum instantaneous operating current to accommodate switching losses during non-zero switching conditions [9, 10]. 600 V, 100 A fast switching IGBTs are used as switching devices in the given application to enable high-frequency switching. A 2 µs dead time is maintained between the switching of IGBTs to allow safe turn-off and to prevent shoot-through of the inverter legs.

Resonant circuit — parallel LC

Induction heating requires high-frequency flux produced by the heating coil to be linked with the workpiece to produce induced currents in the workpiece with an enhanced skin effect. Producing high-frequency flux requires high-frequency current flowing in the heating coil. It is generally difficult to establish high-frequency currents in a coil due to the high inductive reactance naturally exhibited by the coil. Establishing resonance between the inductance of the heating coil and the capacitance of a suitable tank capacitor helps cancel the inductive behavior of the heating coil, enabling easy passage for high-frequency currents through the heating coil.

The resonant circuit for an induction heater could be based on either a parallel LC circuit or a series LC circuit. Figure 5 shows the typical schematic of a parallel resonant induction heating circuit with a capacitor bank ${C}_{2}$ connected in parallel with the heating coil with reactance ${X}_{L}$. Resistance of the steel pipe to be heated ${R}_{L}$ is reflected on the heating coil in such a way that the effective impedance of the heating coil becomes ${{{R}_{L}}^{\mathrm{^{\prime}}}+jX}_{L}$ [1]. Parallel resonant induction heating circuits require the usage of impedance matching/filter inductor ${L}_{m}$ or a current source inverter for enabling efficient power delivery to the heating load [7]. Power delivered to the heating load could also be controlled to some extent by adjusting the reactance value ${X}_{m}$ of the impedance matching inductor, although this method of power control is not recommended.

Fig. 5

Parallel resonant isolated heating circuit

The inverter output voltage and the current for a parallel resonant induction heater are shown in Fig. 6. The main drawback of the parallel resonant induction heater is that even though the tank circuit current or the individual currents through the heating coil or capacitor is sinusoidal, the inverter current is non-sinusoidal due to suppression from the tank circuit near peaks of each cycle, which makes it difficult to transfer power from the inverter to the tank circuit for effective heating of the workpiece. Moreover, the parallel resonant induction heating circuit must operate exactly at its resonant frequency to keep the inverter current minimum. Frequencies higher than the resonant frequency increase the inductive reactance of the heating coil decreasing the coil current and the heat produced on the workpiece, on the other hand, decreases the capacitive reactance, causing an increase in capacitor current, which is undesirable. Frequencies lower than the resonant frequency cause reduction in skin effect, drawing higher currents from the inverter for the same heat produced.

Fig. 6

Inverter output voltage (yellow) and current (pink) of parallel resonant induction heater

With low values of matching inductor ${X}_{m}$, the inverter current becomes increasingly reactive despite the inverter operating at resonant frequency due to the rejecting behavior of the parallel resonant tank circuit [6]. The inverter current becomes almost fully reactive as shown in Fig. 7 when ${X}_{m}$ is close to zero. It can be observed that there is a change in inverter voltage within half cycles due to the high reactive currents charging and discharging the DC link.

Fig. 7

Inverter voltage (yellow) and current (pink) of a parallel resonant induction heater with low matching inductor

Figure 8 shows how the inverter current is shaped by the influence of the sinusoidal voltage of the parallel resonant tank circuit and the inverter switching. The graph showing the inverter current and IGBT gate pulses for the best-adjusted impedance matching inductor is shown in Fig. 9. Higher value of the matching inductor would reduce the amplitude of the inverter current, reducing the output power.

Fig. 8

Resonant current (yellow) and inverter current (pink) of parallel resonant induction heater

Fig. 9

IGBT gate voltage (yellow) and inverter current (pink) of a parallel resonant induction heater with high matching inductor

Resonant circuit — series LC

The resonant circuit can also be designed with the resonant capacitor C₂ in series with the heating coil ${X}_{L}$ as shown in Fig. 10. The significant advantage of a series resonant circuit for induction heater is that it readily accepts power from the inverter without the need of a matching inductor as in the parallel resonant induction heating circuit [11,12,13]. Unlike the parallel resonant induction heater circuit, the inverter operation in a series resonant induction heater can be safely deviated away from resonance frequency with the inverter current reducing. However, the inverter is always controlled “above-resonance” to prevent capacitive switching at frequencies lower than resonant frequency [14].

Fig. 10

Series resonant isolated heating circuit with secondary side resonant capacitor

The secondary circuit carries large resonant currents that flow through the heating coil. It is therefore important that the tank capacitor ${C}_{2}$ and the secondary winding of the transformer are rated for such high current values. The transformer is designed with a ratio of 1:12. Small voltages are usually sufficient to drive a series resonant circuit as the resonant voltages across the heating coil and the capacitor ${C}_{2}$ are in phase opposition with each other. The resonant current in the tuned secondary circuit being naturally sinusoidal [2], forces the inverter current on the primary side also to be sinusoidal in shape as shown in Fig. 11. This enables resonant switching in the inverter side and results in low switching loss at resonance [1, 2, 10].

Fig. 11

Resonant current/inverter current of series resonant induction heater

The capacitance required could be calculated from the expression of resonant frequency as [8]

$${f}_{r}=\frac{1}{2\pi }\sqrt{L{C}_{2}}$$

(4)

The choice of resonant frequency is based on factors such as the type of switching device, material to be heated, and audio frequency range. Ultrafast IGBTs are suitable for switching frequencies up to 100 kHz. However, the frequency should be kept as low as possible to minimize the switching losses. If the electrical resistance of the material to be heated is low, the operating frequency needs to be increased sufficiently high to decrease the skin depth and increase the resistance. This is the reason why heating aluminum or copper requires frequencies from 100 kHz to over 1 MHz. Keeping the resonant frequency below 20 kHz may create audible noise from the system [8]. Based on the above understanding, the operating frequency is selected as 33 kHz.

With a desired resonant frequency of 33 kHz given the inductance of the heating coil 1. 94 µH,

$$\begin{array}{l}C_{\mathit2}\;=\;\frac1{4\mathrm\pi^2\;\times\;33,\;000^2\;\times\;1.94\;\times\;10^{-6}}\;\\C_2\;=\;12.01\;\upmu\mathrm{F}\end{array}$$

(5)

An alternative design could have the resonant capacitor on the primary side of the circuit as shown in Fig. 12 [8,9,10]. It would allow the capacitor to be loaded with low currents from the inverter rather than high currents in the secondary. The transformer ratio would reduce to around 1:2 in order to accommodate the resonant voltages. The capacitor would be subjected to higher resonant voltages as many times as the transformer ratio [6].

Fig. 12

Series resonant isolated heating circuit with primary side resonant capacitor

With the transformer ratio as 1:12, the value of capacitance required on the primary side for the same resonant frequency is calculated as [15]

$${C^{\prime}}_{2}=\frac{{C}_{2}}{{k}^{2}}$$

(6)

where, $k$ is the transformer ratio = 12.

Substituting the value of ${C}_{2}$ from (2),

$$\begin{array}{cc}{C^{\prime}}_{2}=& \frac{12. 01\times {10}^{-6}}{{12}^{2}}\\ {C^{\prime}}_{2}=& 83. 33 nF\end{array}$$

Impedance matching transformer

The impedance matching transformer primarily helps in matching the voltage and current levels of the inverter side and load side. The transformer ratio depends on the type of resonant circuit and other parameters. Ferrite with high core resistance and the ability to handle flux densities up to 1500 Guass was found suitable as the core material for the transformer. Other types of cores such as sendust and iron powder cores, even though capable of handling higher flux densities up to 10,000 to 15,000 Guass, produce a high amount of heat due to core losses. Minimum number of primary turns considering the fundamental component primary winding of the transformer is calculated using its emf equation, to prevent saturation of the transformer core during circuit operation [16].

$${N}_{pri}=\frac{{V}_{dc}\times { 10}^{8}}{4\times f\times {B}_{max}\times{A}_{c}}$$

(7)

where ${V}_{dc}$ is the DC link voltage (V)

$f$ is the inverter frequency (Hz)

${B}_{\mathrm{max}}$ is the maximum flux density (Guass) and

${A}_{c}$ is the area of the magnetic core (cm²).

$$\begin{array}{l}N_{pri}\;=\;\frac{325\;\times\;10^8}{4\;\times\;33\;\times\;10^3\;\times\;1300\;\times\;15.6}\\N_{pri}\;=\;12.14\end{array}$$

Driver power supply isolation

The driver circuit for the IGBTs would normally require a separate power supply as in this application. It would seem convenient to use a switch mode power supply (SMPS). However, it is extremely important to note that these SMPS modules have Class-Y safety capacitors (few nano-farad) linking the AC power supply side to the output for suppressing electromagnetic interference. With the SMPS connected as a power supply to the IGBT driver circuits, a weak capacitive link is formed between any one of the AC supply terminals and the DC link as shown with the capacitor ${C}_{s}$ in Fig. 13, resulting in the malfunction of the controller due to high-frequency switching noise from the IGBTs.

Fig. 13

Power circuit showing the equivalent of Class-Y safety capacitor appearing between the AC supply and the DC link

At no load or lightly loaded conditions, the capacitor ${C}_{s}$ charges up the DC link voltage up to twice the peak value of the AC supply voltage, which would damage the driver ICs and the IGBTs if they are not rated for such high voltages. For 230 V AC supply voltage, the overvoltage on the DC link could reach over 600 V as shown in Fig. 14.

Fig. 14

Graph showing over voltage issue due to capacitive coupling between the AC source and the DC link

The overvoltage issue can be overcome by using truly isolated power supplies or transformer-based rectifier circuits as power supply for the IGBT driver circuits.

Heating power control

Power control in a series resonant induction heater could be achieved in various ways. Controlling the AC input to the rectifier, using a controlled rectifier, and controlling the inverter output through PWM (pulse width modulation) are obvious ways. While the first two methods are does not have any undesirable effect, PWM moves the control away from resonant switching and causes a significant increase in switching losses even at low power levels [7, 11].

Power control can also be accomplished by changing the transformer ratio. However, this method could not be adopted to control power during the circuit operation.

Another interesting method of controlling power is to adjust the operating frequency of the inverter with respect to the resonant frequency as shown in Fig. 15 [15, 17]. The power output reduces when the inverter frequency is moved away from the resonant frequency on either side. With a square wave inverter, it is important that the frequency is adjusted above resonance [17] in order to avoid high $di/dt$ inverter currents due to the capacitive nature of the circuit below the resonant frequency that may result in failure of the switching devices.

Fig. 15

Power–frequency curve of series resonant induction heater showing the region of power control

In the given application, the inverter frequency is adjusted to 37 kHz to set the maximum power level at 3. 5 kW during operation. It should however be noted that due to the inverter operation slightly deviating from resonance, there would be an increase in switching losses [11], accompanied with a reduction in conduction losses due to lower inverter currents, making the overall increase in the heat produced by the IGBTs insignificant.

Results and discussion

System implementation and experimental findings

The system is implemented in the operational environment of the servo valve test setup as per the block diagram shown in Fig. 2. The power circuit with AC voltage controller, diode bridge rectifier, and water-cooled IGBT inverter with polyester DC link capacitors is shown in Fig. 16. The controller is isolated from the driver and power circuit through high-speed opto-couplers to ensure safety with resulting IGBT gate pulses as shown in Fig. 17. The charging and discharging behaviors of the gate capacitance of the IGBTs, and the 2. 5 µs dead time could also be observed. An alternative design of the IGBT driver circuit was experimented with pulse transformers to provide isolated gate pulses for multiple IGBT, but was not considered for the implementation due to severe distortions observed in the gate pulses due to inductive coupling, which could be seen in Fig. 18.

Fig. 16

Power circuit of the induction heating system showing the AC voltage controller, rectifier, and the IGBT inverter

Fig. 17

IGBT Gate pulses obtained from driver circuit based on high-speed opto-coupler

Fig. 18

IGBT Gate pulses obtained from pulse-transformer-based driver circuit

Figure 19 shows the power circuits with the accessories such as auxiliary power supplies and the cooling system for IGBTs and the heating coil. The high-frequency transformer should preferably be placed as close as possible to the inverter output to avoid stray inductance of the cables affecting the performance of the circuit. A pair of twisted cables is used for an unavoidable distance of 2 m between the inverter output and the transformer, as both couldn’t be brought any closer. Twisted cables help break large inductance loops into smaller ones and bring bipolar currents close to each other to help reduce the effect of stray inductances.

Fig. 19

Power circuit of the induction heating system with accessories

The impedance matching transformer, capacitor bank, and heating coil are arranged as shown in Fig. 20. The heating coil and the secondary winding of the transformer are water cooled to help the removal of heat produced by the large magnitudes of the resonant currents in the secondary circuit. The 12 µF capacitor bank is realized with individual 1 µF high current polyester 1 µF capacitors connected in parallel. In this case, the individual capacitors could be connected in series to obtain the capacitance of 83. 33 nF and connected in the primary side of the circuit for the same resonant frequency. It should be noted that the one-to-one relationship of capacitance values between parallel series connections is because of the transformer ratio is 1:12.

Fig. 20

Heating circuit with inductive heating coil, capacitor bank, and the impedance matching transformer

The arrangement of the inductive heating coil in the servo valve test setup can be seen in Fig. 21. Electrical isolation due to impedance matching transformer proves induction heating to be much safer than resistive heating in addition to the other benefits such as isolation from high pressure and clean heating.

Fig. 21

Servo valve test setup with induction heating arrangement

The actual temperature profile of hours 176 to 180 out of 300 continuous hours of servo valve testing is shown in Fig. 22 and that of hours 293 to 300 is shown in Fig. 23. The 10-min temperature ramp from 45 °C to 90 °C, 90 °C to 45 °C and the steady temperature level of 90 °C for 35 min are achieved as per the desired temperature profile.

Fig. 22

Actual temperature profile during hours 176 to 180 of the servo valve test

Fig. 23

Actual temperature profile during hours 293 to 300 of the servo valve test

The induction heater handled the desired power rating of 4 kW with low heat losses due to the near-resonant switching. Smooth control of temperature is achieved due to the efficient operation of the induction heating system with closed-loop control. A highly reliable heating arrangement for the high-pressure testing application at 84 bar with inline heated fuel is developed based on a clear understanding of the functional differences between the various configurations of the induction heating system and the challenges in designing the individual component of the system.

Efficiency analysis

The efficiency analysis of the induction heating system helps us determine how effectively the system converts electrical energy into heat and how efficiently it transfers that heat to the target substance, in this case, the aviation fuel. Various factors influence the overall efficiency of the induction heating system, such as the design of the resonant tank circuit, the selection and operation of the IGBT inverter, and the cooling arrangements for the power components. Additionally, the impedance matching transformer and capacitor bank configuration can impact the energy transfer efficiency. By carefully analyzing these factors and quantifying the losses incurred at different stages of the heating process, we can gain insights into the system’s energy efficiency and identify areas for improvement. The efficiency analysis provides insights into the magnitude of power losses occurring in various components of the system, including the power circuit, driver circuit, and auxiliary components like the ferrite core transformer, heating coils, and cables.

Extensive simulations were carried out in LTspice to assess the power losses occurring in the various components of the power conversion circuits. The power electronic converter exhibited efficiencies between 94 to 96% in the simulations. Most of the power losses are attributed to the switching losses and conduction losses in high-frequency electronic switching devices and the conduction losses in the bridge rectifier. However, the actual system exhibited additional power loss in the distributed resistance of the electrical cables and due to heat radiation near the outer surface of the workpiece through its insulation layer of glass wool. These observations were further validated through experimental measurements of efficiency.

The output power that is transferred to the fuel as heat is calculated as [18]

$${P}_{out}= d\times q\times C\times \Delta T$$

(8)

where $d$ is the density of the Jet A1 fuel (840 kg/m³),

$q$ is the flow rate of the fuel inside the heating pipe (1 lpm = 16.667 × 10⁻⁶ m³/s),

$C$ is the specific heat of the fuel (2.2 kJ/kg°K), and

$\Delta T$ is the difference in temperature after and before heating (90–45 °C = 45 °C).

$$\begin{array}{l}{P}_{out}=840 \times 2.2\times {10}^{3} \times 16.667 \times {10}^{-6} \times 45\\{P}_{out}= 1.386\mathrm{\, kW}\end{array}$$

The input power to the induction heating system is given by

$${P}_{in}= k\;\times{V}_{rms}\times {I}_{rms}$$

(9)

where $k$ is the duty ratio of the AC voltage controller,

${V}_{rms}$ is the rms value of input voltage (V), and

${I}_{rms}$ is the rms value of input current (A).

$$\begin{array}{l}{P}_{in}= 0.45\times 230\times 14.6\\{P}_{in}= 1.511\mathrm{\, kW}\end{array}$$

The efficiency of the induction heating system is given by

$$\begin{aligned}\eta =\ &\frac{{P}_{out}}{{P}_{in}}\\ \eta =\ & \frac{1.386}{1.511}\\ \eta =\ & 0.917\mathrm{\, or\, }91.7\mathrm{\%}\end{aligned}$$

(10)

Based on the calculated efficiency of 91.7% for the heating system, it is evident that there is a drop in efficiency compared to the power electronic converter’s efficiency of 94 to 96%. This decrease in efficiency can be attributed to various factors, including heat radiation from the fuel pipe, which results in heat loss rather than efficient heat transfer to the fuel. However, it should be noted that these efficiency values are much higher than the practical efficiencies of other heat exchanger-based systems which lie around 60–70% or even less. Further analysis and optimization efforts can focus on mitigating heat losses and improving heat transfer mechanisms to achieve higher efficiencies in future iterations of the induction heating system.

Comparison with alternative designs

In comparison to the commonly used high-pressure heat exchanger-based system for fuel heating, our proposed inline induction heating system offers several significant advantages. The high-pressure heat exchanger-based system utilizes thermic fluid and water for heating and cooling the fuel, respectively, resulting in energy losses due to large thermal inertia, increased complexity, and higher costs. Dedicated systems are required for heating the thermic fluid and cooling the water in a heat exchanger-based system, which makes the system much more complex. Moreover, high-pressure heat exchangers are typically made of steel, which exhibits poorer heat transfer response compared to low-pressure aluminum heat exchangers. The system also requires the use of flow control valves for the thermic fluid and cooling water, further contributing to its complexity and cost.

In contrast, our inline induction heating system provides localized heating, rapid response, efficient cooling due to low-volume heating, fewer heat transfer components, precise control, and overall cost savings. The design cost of our inline induction heating system was approximately ₹ 15,000, significantly lower than the estimated cost of approximately ₹ 10 lakh for the heat exchanger-based system.

Commercially available induction heating systems have been widely used for various applications. These systems employ induction heating as the primary mechanism, providing efficient and controlled heating. However, when it comes to high-pressure inline heating, these off-the-shelf systems may not be suitable without certain modifications. Typically, these systems require adjustments to the resonant components and the incorporation of closed-loop control to meet the specific requirements of inline heating. Additionally, the control capabilities of these systems are often limited, relying on manual control options. Incorporating automatic control may necessitate further modifications. Despite their moderate energy efficiency, these commercially available systems lack optimization for specific applications, resulting in moderate material efficiency. The complexity of repurposing these systems for high-pressure inline heating is also a factor to consider. Moreover, the cost of such systems tends to be high, ranging from ₹ 3–5 lakh, and their maintenance can incur significant service costs. Safety is another concern, as the heating coils may not be isolated from the mains, and the system behavior may be unpredictable.

The following table provides a comprehensive comparison between our proposed inline induction heating system, the heat exchanger-based system, and commercial induction heating systems, highlighting their key attributes and performance parameters (Table 1 ).

Table 1 Proposed induction heating system vs. heat exchanger-based system

The comparison between our proposed inline induction heating system, the heat exchanger-based system, and commercial induction heating systems reveals distinct advantages of our proposed solution for high-pressure fuel heating applications. The inline induction heating system offers localized heating, almost instantaneous response, precise control, high energy and material efficiency, low complexity, cost-effectiveness, minimal maintenance requirements, and enhanced safety measures. In contrast, the heat exchanger-based system exhibits limitations in terms of slower response, increased energy losses, higher complexity, and maintenance demands. Commercial induction heating systems, while not readily available for high-pressure inline heating, require modifications and lack customization options. Based on these findings, our proposed inline induction heating system emerges as a superior choice, offering a compelling alternative for high-pressure fuel heating applications with its exceptional performance, efficiency, and cost-effectiveness.

Conclusions

The present study aimed to develop an efficient and reliable induction heating system for high-pressure fuel heating applications. Extensive research was conducted on various configurations of induction heater circuits, considering factors such as power efficiency, system complexity, and heat transfer mechanisms. Based on this analysis, the high-pressure fuel inline heating system with a series resonant induction heater with isolation using an impedance matching transformer was selected as the most suitable configuration.

The proposed induction heating system offers several distinct advantages over alternative designs, particularly the commonly used high-pressure heat exchanger-based system. By utilizing localized heating and precise control, our system provides almost instantaneous response and high energy efficiency. Furthermore, the design cost of the proposed inline induction heating system is significantly lower than that of the heat exchanger-based system, making it a cost-effective solution for high-pressure fuel heating.

This study provides a comprehensive analysis of induction heater circuits, considering various factors such as power conversion efficiency, impedance matching, and driver power supply isolation. The findings not only contribute to the understanding of different configurations but also provide practical insights for designing efficient and reliable induction heating systems for specific applications.

The developed induction heating system demonstrates exceptional performance, energy efficiency, and cost-effectiveness, making it a compelling alternative for high-pressure fuel heating applications. This research contributes valuable knowledge on induction heater circuits and provides a practical solution for high-pressure fuel heating. The findings pave the way for further optimization and future advancements in induction heating technology, ultimately benefiting industries that require efficient and reliable heating processes.