Laser ignition system on CNG Engine


Master's Thesis, 2013
67 Pages

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TABLE OF CONTENTS

Chapter 1 Introduction
1.1 LITERATURE SURVEY
1.2 Fundamentals of Ignition
1.3 Why Laser Ignition?
1.4 Goals and Objectives
1.5 Natural Gas-Air Ignition Experimental Study

Chapter 2 Fundamental of Ignition
2.1 Laser Ignition System
2.2 Conventional Ignition System
2.3 Operational Procedure
2.4 Results and Discussion
2.4.1 Test Matrix
2.4.2 Ignition Limits
2.4.3 Minimum Required Energy Scans
2.4.4 Ignition Delays and Rates of Pressure Rise
2.4.5 Conclusions

Chapter 3 Design of LIS Components
3.1 Laser System
3.2 Laser Plugs
3.3 High-Power Optical Multiplexer
3.3.1 Electro-Optic Modulator (Pockels Cell)
3.3.2 Rotating Mirror
3.3.3 Flip-flop
3.4 Fiber-Optic Delivery
3.4.1 Solid Core Fibers
3.4.2 Advanced Air-Core Fibers
3.5 Electronic Interface
3.6 Results and Conclusions,

Chapter 4 Engine Testing Experimental Setup and Procedure
4.1 Engine Test Cell
4.2 Experimental Setup
4.2.1 Single-Cylinder Engine
4.2.2 Open-Path Laser Ignition Setup
4.2.3 Fiber-Coupled Laser Ignition Setup
4.3 Combustion Quality Measurement
4.4 Exhaust Gas Analysis
4.4.1 NOx Analysis
4.4.2 CO and CO2 analysis
4.4.3 Total Hydrocarbon Analysis
4.5 Results and Discussion
4.5.1 Full Load Comparison (15 bar BMEP)
4.5.2 Part Load Comparison (10 bar BMEP)
4.5.3 Fiber-Coupled Laser Ignition Results
4.6 Conclusions

5. References

CERTIFICATE

Certified that this report titled “LASER IGNITION SYSTEM ON CNG ENGINE, for the phase-II of the project, is a bona fide work of Mr. VIKAS SHARMA (11PEIC1004), who carried out the work under my supervision, for the partial fulfillment of the requirements for the award of the degree of Master of Technology in Internal Combustion Engine . Certified further that to the best of my knowledge and belief, the work reported herein does not form part of any other thesis or dissertation on the basis of which a degree or an award was conferred on an earlier occasion.

Chapter Introduction

1.1 LITERATURE SURVEY

Economic as well as environmental constraints demand a further reduction in the fuel consumption and the exhaust emissions of motor vehicles. At the moment, direct injected fuel engines show the highest potential in reducing fuel consumption and exhaust emissions. Unfortunately, conventional spark plug ignition shows a major disadvantage with modern spray-guided combustion processes since the ignition location cannot be chosen optimally. It is important that the spark plug electrodes are not hit by the injected fuel because otherwise severe damage will occur. Additionally, the spark plug electrodes can influence the gas flow inside the combustion chamber.

On the other hand, a laser has been discussed widely as one of the promising alternatives for an ignition source of the next generation of efficient internal combustion engines (Hickling & Smith, 1974; Dale et al., 1997; Phuoc, 2006). Laser ignition can change the concept of ignition innovatively and has many advantages over conventional electric spark plug ignition.

It is well known that short and intensive laser pulses are able to produce an”optical breakdown” in air. Necessary intensities are in the range between 1010 . . . 1011W/cm2. At such intensities, gas molecules are dissociated and ionized within the vicinity of the focal spot of a laser beam and hot plasma is generated. This plasma is heated by the incoming laser beam and a strong shock wave occurs. The expanding hot plasma can be used for the ignition of a combustible material. Other laser ignition methods, like thermal ignition of a combustible due to heating of a target or resonant absorption which generates radicals are not able to fulfill the requirements on a well defined ignition location or time and will not be discussed further. In the past, this optical breakdown has been used for ignition of gas mixtures many times. In most cases, only slow combustion processes have been investigated. This article will present some basics of laser ignition together with results achieved by operating a laser ignition system on an internal combustion engine for a long period of time. Basics of fast combustion processes will be discussed briefly. Figure 1 shows the schematics of combustion engines ignited by (a) an electric spark plug and a laser (b), (c).

Using a laser, the ignition plasma may be located anywhere within the combustion chamber because laser ignition doesn’t need electrodes. Optimal positioning of ignition apart from the cold cylinder wall allows the combustion flame front to expand rapidly and uniformly in the chamber and thus increases the efficiency as seen in (b).

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Fig.1.1. Schematics of the combustion engines ignited by (a) a spark plug and (b), (c) a laser. (c) Shows multipoint ignition.

In addition, laser ignition has great potential for simultaneous, spatial multipoint ignition within a chamber as shown in (c). This shortens combustion time dramatically and improves the output and efficiency of engines effectively. Further a laser can ignite leaner or high pressure mixtures that are difficult to be ignited by a conventional electric spark plug. A laser igniter is also expected to have a longer lifetime than a spark plug due to the absence of electrodes.

1.2 Fundamentals of Ignition

In a typical spark plug, successful sparking is achieved when the potential drop across the spark gap exceeds the dielectric breakdown threshold of typical gases. The sparking potential across the electrodes is given by Paschen’s law:

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Where, p is the pressure of the gas and d is the spark gap. The breakdown voltage, Vb, exhibits a linear dependence on the product pd. The electrode shape and material are also found to have significant influence on the spark ignition process. Once gas breakdown occurs and a plasma kernel is established, energy transfer occurs mainly through diffusion on the surface to the surrounding gas. Whether such a diffusion process results in a successful combustion flame front depends upon the kernel energy exceeding Minimum Ignition Energy (MIE), the kernel size exceeding a certain size, turbulence and gas speed. In practice, factors influencing successful spark creation far outweigh those influencing its transformation into a flame front, and it is normally assumed that once a spark is created the mixture is successfully burned.

1.3 Why Laser Ignition?

In present turbocharged of the lean-burn natural gas engines, Capacitance Discharge Ignition (CDI) systems are used as schematically shown in Figure 4. Though these systems are rated at 100-150 millijoules (mJ) per strike, after thermal losses typically 40-60 mJ is transmitted to the spark kernel at rates of voltage rise of 500 Volts per millisecond (V/μs). In CDI systems, energy stored in a high-voltage capacitor (at ~ 175 Volts Direct Current (VDC) is discharged through a high-voltage coil resulting in voltages in excess of 28 kilovolts direct current (kVDC) across the spark plug gaps.

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Figure 1.2. Schematic of a Capacitance Discharge Ignition (CDI) System.

With a push towards lean-engine operation, with a concomitant requirement to maintain engine specific power, the intake air pressure is increased. Lean operation along with high intake air pressure results in very high charge densities at the time of ignition. Such high gas densities necessitate spark gap voltages in excess of 40 kilovolts (kV) that cannot be achieved using current CDI systems. This often leads to increased misfiring with subsequent loss of fuel efficiency and increased unburned hydrocarbon (UHC) emissions. Higher UHC emissions are essentially volatile organic compounds (VOC), which are currently regulated in California. To address these problems, various research organizations have been exploring alternate ways to achieve ignition. Among these alternate methods, laser ignition proves attractive as it offers the following performance benefits:

- Successful ignition of mixtures at high pressures ensures reduced occurrence of misfire, and consequently improved fuel efficiency and lower UHC emissions,
- Potentially lower maintenance as the requirement to maintain a reasonable spark gap is eliminated,
- Extension of lean operating limits, thereby enabling lower NOx emissions,
- Shorter ignition delays and enhanced combustion rates, which allow retarded
- timings thereby reducing NOx emissions, and
- Location of ignition kernel away from the walls, thereby enhancing overall efficiency
- due to reduced heat loss to the cylinder head.

1.4 Goals and Objectives

Laser ignition can overcome the ignition problems in lean-burn natural gas engines and further has the potential to improve engine efficiency and lower emissions. The overall benefits due to laser ignition can be summarized as:

- Improved overall efficiency,
- Reduced fuel consumption,
- Lower NOx and unburned hydrocarbon (UHC) emissions,
- Enhanced power density, and
- Reduced overall maintenance requirements.

The overall goals of the proposed LIS system are:

- Meet or exceed the current and future California emissions requirements and have other desirable environmental attributes.
- Improve fuel-to-electricity conversion efficiency.
- Lower capital costs, installation costs, operation and maintenance cost, and life cycle costs.
- Enhance reliability, maintainability, durability and usability.
- Possess multi-fuel use capabilities, such as with sewer gas, landfill gas etc.

1.5 Natural Gas-Air Ignition Experimental Study

With the renewed interest in laser ignition, there have been quite a few past studies evaluating laser-based ignition in static chambers. As shown in Figure 6, such studies have shown that lasers enable ignition of mixtures at pressures higher than those that can be ignited by conventional coil based Capacitance Discharge Ignition (CDI) systems. However, no significant extension of lean-ignition limit was found with laser ignition. The lean ignition limits for both modes of ignition appeared to coincide at ϕ equal to 0.67.

With most of the lean-burn engines operated close to the intersection of lean ignition limit and self-ignition limit, such a spread in data warrants a systematic study under in cylinder like conditions. To attain this goal, one needs to perform ignition tests in a Rapid Compression Machine (RCM) that simulates typical natural gas engine conditions. As most of the lean burn engines are operated close to the intersection of ignition limit and knock limit, the self-ignition limit needs to be determined as well. Also, mapping of the minimum laser energies required for successful ignition under different mixture conditions would assist in the development of LIS.

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Figure 1.3. Ignition Limits of Methane-Air Mixtures Established in a Static Chamber (Initial Mixture Temperature ~ 22° C).

Chapter 2 Fundamental of Ignition

2.1 Laser Ignition System

The laser ignition system is schematically shown in Figure 2.1. The beam output of a frequency doubled Neodymium: Yttrium-Aluminum-Garnet (Nd:YAG) laser was routed through a combination of half-wave plate and polarizer to vary the laser power. A beam-splitter and a power meter allowed monitoring of the laser power. A fast-shutter with 3 millisecond time response allowed incidence of a single pulse from the laser pulse train. All of the timing was controlled by the NI-Field Point system.

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Figure 2.1 . Schematic of the Optical Arrangement in the Laser Ignition System.

2.2 Conventional Ignition System

A CDI system, modified for the present tests. This system, when activated remotely by a 5 volt (V) pulse supplies ignition energy to the spark plug placed on the wall of the combustion chamber. The arrangement was such that the spark was supplied 30 ms following the end of piston stroke. Such a delay was necessary to allow the locking mechanism to engage completely before the initiation of combustion. Figure 2.2. shows a picture of this ignition system.

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Figure 2.2. Conventional Ignition System Cart.

2.3 Operational Procedure

In a typical experiment, the RCM pistons were retracted and held in the retracted position. Subsequently, a gas mixture of the required equivalence ratio and initial pressure, P1, was established in the compression chambers. After allowing 5 minutes for the gases to mix, the pneumatic chambers were pressurized with compressed air at 20.7 bar supplied by a 150 Liter air tank. The pistons were released by activating the appropriate valve sequencing. A photo detector sensing the piston position provided the necessary signal for sequencing the laser pulse or the conventional ignition spark. To allow for locking of the pistons in the compressed position and thereby avoid a piston bounce back at the commencement of combustion, ignition was initiated 30 ms following the end of compression stroke. Also, one second following the end of compression stroke, the exhaust valve was opened. Subsequently, the pistons were retracted and the compression and combustion chambers were purged to prepare for the next experimental run. A typical test run required 20 to 30 minutes for execution.

2.4 Results and Discussion

2.4.1 Test Matrix

With the above setup, tests were performed while varying the initial pressure of the mixture, P1, and the equivalence ratio, ϕ. The mixtures established in the RCM were limited to 1.0 > ϕ >0.4, 3.0 > P1 >1.0 bar and initial temperature, T1 = 298 K. A typical pressure trace obtained during one such test run is shown in Figure 2.3. As shown in Figure 2.3, when the pistons are released, the gaseous mixture is isentropically compressed to P2. As the ignition (sparking) event is sequenced 20 to 30 ms following end of compression, there is a small pressure drop resulting from heat transfer to the combustion chamber walls, (P2-P2’).

Following an ignition delay after the incidence of spark, the pressure rises steeply to P3 due to combustion. Subsequent pressure drop is primarily due to condensation of water vapour and heat transfer to walls of the chamber. From thermodynamics, one has the relations.

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where, CR is the compression ratio and γ is the ratio of specific heat at constant pressure (Cp) to specific heat at constant volume (Cv), which equals 1.4 for air.

A regression analysis performed on the measured values of P2 assuming γ as 1.4 showed that the compression ratio for the present RCM is 10.0 as opposed to 12.0, as shown in Figure 2.4, it was observed that the measured P3 values were on an average 83% of the calculated P3 values, with leaner mixtures exhibiting lower values. Also, it ought to be noted that in the present tests, for all mixture conditions, the temperature at the time of ignition, T2 was about 765 K (per Equation 2.2).

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Figure 2.3. Typical Pressure Traces from RCM Operation; P1 = 1 bar, φ = 0.7

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Figure 2.4. Measured Versus Calculated Peak Combustion Pressures for Various Methane-Air Mixtures, 1.0 < P1 <3 bar and 0.5 < ϕ < 1.0

2.4.2 Ignition Limits

With the above setup, tests were conducted by varying the equivalence ratio,ϕ, between 0.5 and 1.0, and by varying the initial pressure of the mixture, P1, between 1.0 and 3.0 bar. For each condition, typically 5 or 6 test runs were executed while varying the laser pulse energy between the maximum pulse energy of about 75 mJ and minimum pulse energy of 5 mJ. With each test the window within which the threshold energy was present was halved until the final threshold value was determined within an accuracy of ± 2.25 mJ/pulse.

The ignition boundaries determined through such tests expressed as a function of pressure at the time of ignition, P2’, are shown in Figure 2.5. It is prominently noticed that self ignition dominates for mixtures with P2’ greater than 63 bar. Also, it is noticed that the lean ignition limit while using the CDI system is ϕ at 0.6. On the other hand, by using laser ignition this could be extended all the way to the flammability limit of ϕ of about 0.5. Such extensions are of significance as the current lean burn engines are operated at the intersection of self ignition limits and lean ignition limits

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Figure 2.5. Ignition Boundaries Determined by Using an RCM.

2.4.3 Minimum Required Energy Scans

For the ignition boundaries established above, Minimum Required Energy (MRE) values for successful laser ignition were determined for various mixture conditions. Values determined for a lens focal length, f of 13 millimeters and laser beam quality of M2 ≤ 5 are shown in Figure 2.6. It is observed that except for ϕ of 1.0, MRE values decreased with increase in pressure finally resulting in self-ignition. The values along a cross-section of P2 of 37.7 bar are shown in Figure 2.7. For a variation of the equivalence ratio, it is noticed that a minima exists at ϕ of 0.85 for these methane-air mixtures. Also it is noticed that there is a sharp rise in the MRE values for mixtures leaner than ϕ of 0.7. Such a trend was also noticed at other pressures.

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Figure 2.6. Minimum Required Laser Energies (MRE) for a Lens Focal Length f = 13 mm and Laser Beam Quality of M2 ≤ 5.

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Figure 2.7. Minimum Required Laser Energies for P237.7 bar.

2.4.4 Ignition Delays and Rates of Pressure Rise

Lean natural gas-air mixtures are characterized by slow flame velocities and longer ignition delays that are of concern for lean burn engine operation. Previous studies in 1-cylinder engines have shown that laser ignition results in smaller ignition delays and faster combustion. As shown in Figure 2.8, similar tests performed in the RCM showed that ignition delay increased with lean operation for both CDI ignition as well as laser ignition. However, the benefits in terms of smaller ignition delays were only pronounced under lean conditions and at ϕ of 1.0. Similar comparison of the rate of pressure rise is shown in Figure 2.9. Again it was found that faster combustion rates occur for lean operation and for ϕ of 1.0 by the use of laser ignition. The reversal of trends for 0.75 < ϕ < 0.95 requires further investigation.

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Figure 2.8. Ignition Delays for CDI and Laser Ignition.

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Figure 2.9.Rates of Pressure Rise for CDI and Laser Ignition.

2.4.5 Conclusions

Through ignition tests conducted in a Rapid Compression Machine that simulated incylinder conditions of a lean-burn compressed natural gas engine, it was observed that laser ignition extends the lean operating limit of methane-air mixtures to the flammability limit (ϕ of 0.5) whereas conventional CDI ignition on an average is limited to mixtures richer than ϕ of 0.6.

Minimum Required Energies for successful laser ignition exhibited a sharp increase followed by a plateau region for methane-air mixtures leaner than ϕ of 0.7. Such a trend shows that a laser ignition system developed to operate at ϕ of 0.65 will successfully operate under all other possible operating conditions. To reap the true benefits of laser ignition, the required hardware components for the LIS need to be developed.

Chapter 3 Design of LIS Components

From an initial survey of possible configurations for a laser-based ignition system, two promising concepts were identified:

(i) The laser-per-cylinder concept, and
(ii) The multiplexed laser concept.

In the laser-per-cylinder concept schematically shown in Figure 3.1, a miniature laser is built directly over the cylinder head. The cost of such a configuration is likely to be very high, as a single laser is required for each cylinder. Also, thermal management in the laser system becomes an issue as the cylinder head temperatures could be as high as 130o Celsius (C). However, this configuration does not require the use of a high-power delivery system (e.g., a fiber). As part of the LIS consortium, National Energy Technology Laboratory (NETL) has pursued the development of a laser-per-cylinder ignition system, details of which are provided in Reference.

Alternately, as shown in Figure 3.2, the output of a single laser can be distributed among the various cylinders of a multi-cylinder engine. This configuration benefits from the cost savings that result from the use of a single laser as lasers are the most expensive components of LIS. Also, the laser system is isolated from the heat and vibration of the engine. However, this system requires the development of associated high-power components including laser plugs, optical multiplexer and fiber delivery system.

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Figure 3.1. Schematic of the Laser-Per-Cylinder Concept.

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Figure 3.2. Schematic of the Multiplexed Laser Concept.

The specific objectives of this task are described below:

1. Design and develop laser plugs with the following specifications.

i. Have the same thread size as a conventional spark plug, i.e., M18 x 1.5,
ii. Provide pressure sealing to 3,000 pounds per square inch (psi).
iii. Withstand temperatures as high as 2,400° C.
iv. Minimize first and second surface reflections.
v. Minimize overall laser energy requirements.
vi. Provide sufficient reliability while transmitting laser energies of about 60mj/puls.
vii. Be self-cleaning of any deposits.
viii. Facilitate coupling to fiber optic transmission.

2. Develop fiber-optic systems that are able to transmit the required laser energies for LIS operation. In the process, evaluate the performance of the following technologies, among others, for transmissivity, flexibility, and ease of connectivity.

i. Conventional silica core of 1 to 2 millimeter(s) diameter fibers with appropriate cladding.
ii. Fibers with tapered ends.
iii. Hollow fiber systems.

3. Develop an optical multiplexer capable of distributing the laser output among various cylinders of a natural gas engine, while minimizing transmission losses and facilitating the required ignition timing advance or retard. This can be performed by evaluating the following candidate technologies among others.

i. Fiber-optic telecommunication multiplexer.
ii. Rotating grating-type indexing system.

4. Develop or select a laser system that can,

i. Provide the required laser energies.
ii. Operate with minimal maintenance.
iii. Have a small foot print.
iv. Provide the aforementioned features at a low cost.

5. Develop an electronic interface in collaboration with Altronic, Inc., which integrates the functions of laser head, Electronic Control Unit (ECU), and indexer into one single unit. This electronic interface should enhance ease of installation, improve durability, satisfy the safety requirements of natural gas engines, have a small footprint, be lightweight, and facilitate manufacturing in large quantities.

In addition to the data gathered in the earlier, guidance in design was also derived from performance requirements of the ignition system as specified by engine manufacturers, as given in Table 1.

Table 1. Performance Requirements of an Advanced Ignition System (Courtesy: Caterpillar, Cummins and Waukesha )

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3.1 Laser System

From fundamental physics, it is well-known that when a high-power coherent laser beam is focused, multiple photons are absorbed resulting in the generation of free electrons. These electrons are further accelerated by the field gradients and generate more electrons and ions through an inverse Bremsstrahlung process. The plasma so generated serves as the ignition kernel for the combustible mixture surrounding it. The energy transfer to the surrounding gas mixture is primarily through diffusion, and depending on its magnitude may or may not result in successful combustion. Overall the process is not wavelength specific and is termed “non-resonant multi-photon ionization.” An initial literature survey showed that Nd:YAG lasers are ideally suited for laser ignition applications. Based on the configuration of the laser, the output of an Nd:YAG laser can have wavelengths of 1064 nanometers (nm), 532 nm or 266 nm. The focal spot diameter of a collimated laser beam focused by a lens is given by ...

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where,

wo = focal spot diameter for ideal Gaussian mode laser (microns),

Wo = focal spot diameter for a typical multi-mode laser (microns),

M = is the mode quality,

λ = laser wavelength (microns),

f = lens focal length (cm), and

D = laser beam diameter (cm).

As evident from Equation 4, smaller wavelengths result in smaller focal spot diameters. As a result, the use of smaller wavelengths is desirable to achieve a laser flux density of about 1012 Watts per square centimeter (W/cm2), which is required for sparking. With the harmonic generation process being at best 50% efficient, theoretically speaking, a factor of two advantage is achieved by going to a higher harmonic. However, a compromise is necessary as the system complexity increases. For the present purpose, a wavelength of 532 nm was chosen. From a functional standpoint, it was identified that the laser needs to have the following operational characteristics:

- Wavelength = 532 or 1064 nm,
- pulse width < 7 ns,
- repetition rate = 90 Hz (equivalent to the firing requirements of a 6-cylinder, 4-stroke engine operating at 1800 rpm),
- laser energy per pulse = 65 mJ/pulse,
- Base quality, 1.0 < M2 < 5.0,
- Operating environment temperature = 10 – 40°C,
- Compact laser head with low power requirement, and
- Lifetime > 109 shots or approximately 3000 hours for a 6-cylinder engine.

A survey of current laser technology showed four possible candidates: (i) Passively or actively Q-switched Nd:YAG lasers, (ii) Diode pumped solid state lasers, (iii) Disk lasers, and (iv) Fiber lasers. While all of these could provide the required peak laser powers to achieve sparking, Nd:YAG lasers cannot meet the maintenance requirements because their flashlamps have lifetimes in the range of 10-30 × 106 shots. Also thermal management of Nd:YAG lasers requires bulky external cooling. The cooling requirements are somewhat mitigated in the case of disk lasers and fibers lasers due to the large surface area to volume ratio of the lasing media. On account of the rapid advancements in fiber laser technology, it is expected that disk lasers would become obsolete in the near future [16]. However, as of 2005, both fiber lasers and disk lasers were cost-prohibitive for the current application.

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Figure 3.3. A Commercially-Available Diode Pumped Solid State Laser (DPSSL) Model: Centurion, Manufacturer: Big Sky Laser, Inc., 100 Hz, 45 mJ/pulse.

Diode Pumped Solid State Lasers (DPSSL), on the other hand, are ideally suited for the present purpose. With the cost of laser diodes continually decreasing and their power continually increasing, DPSSL are anticipated to make many new applications possible. Most attractive to the current application is the fact that the only component that requires maintenance, the laser diode, has a lifetime greater than 109 shots, which for a typical ARICE engine can translate to 2.5 years. An example of a commercially available DPSSL is shown in Figure 3.3. The performance data of DPSSL from three different manufacturers is tabulated in Table 2 below. The MK-88 laser system from Kigre, has the smallest footprint and has already been shown to achieve sparking in the lab using a 13 millimeter (mm) focal length lens. However, the repetition rate of this laser makes it suitable only for the laser-percylinder application shown in Figure 18. On the other hand, the systems from Big Sky laser and JMAR are suitable for the present application. Table 2 provides a snapshot of the DPSSL technology available as of 2013. This technology is making rapid progress with a very favorable cost and performance trajectory amenable to the laser ignition application. Based on these considerations, it was decided to postpone the selection of a specific laser for the present application. Consequently, all of the Demonstrations in the present project are performed using compact Nd:YAG lasers. Except for longevity of the pump source, the performance of these lasers will mimic that of DPSSL in all other respects,

Table 2 Performance Specifications of Some Commercially-Available Pulsed DPSSL

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Information not available

3.2 Laser Plugs

Laser plugs are elements that introduce an optical window so that laser radiation can be focused to create a spark inside the cylinder. Additional requirements for their performance are given below: Foot print similar to that of standard 18 mm spark plug, Pressure rating ~ 300 bar, Temperature rating ~ 3,000 K on elements exposed to combustion, and should be self-cleaning of carbon and oil deposits. After a couple of iterations, a two-lens design was found to be appropriate for LIS. As shown in Figure 21, the output of a fiber optic cable is collimated using a plano-convex lens. The collimated output is refocused inside the cylinder using a sapphire lens of 13 mm back focal length. Such an arrangement allowed refocusing to a spot size of 240 micrometers (μm). Ray tracing iterations performed using ZEMAX software showed that the use of multi-element lenses does not reduce the final spot size any further.

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Figure 3.4. Ray Propagation Scheme Inside a Laser Plug.

In the arrangement shown in Figure 3.4, the thickness of the plano-convex sapphire lens was chosen to withstand pressures up to 300 bar. With such a thick lens, it is very important that the curved side of the lens be pointed towards the laser so as to avoid internal reflections that will lead to internal cracking of the lens. Such a lens was also found to withstand typical in-cylinder combustion temperatures. The physical arrangement of the aforementioned lens configuration is shown in Figure 3.5. As noticed in this arrangement, the laser plug consisted of a top part and a bottom part separated by an aluminum spacer. The bottom part carried the necessary external threads to fasten the entire laser plug assembly into the cylinder head. A copper crush gasket in the bottom part provided the necessary sealing to prevent leakage of combustion gases beyond the sapphire lens. The top part carried a lens tube, which in turn, housed the collimating plano-convex lens. A fiber optic connector separated from the collimating lens allowed coupling of the incoming laser radiation. Through tests performed by incorporating this laser plug in a 4 kilowatts (kW) natural gas engine, it was observed that the laser flux density on the exit face of the sapphire lens was critical. At low laser flux densities distinct carbon deposits were visible on the lens. Above a certain threshold, the laser radiation ablated such deposits and the lens was found to be self-cleaning. Caution needs to be exercised to keep the laser flux densities well below the material damage threshold of 5 giga watts per square centimeter (GW/cm2).

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Figure 3.5. Schematic of a Two-Lens Laser Plug.

3.3 High-Power Optical Multiplexer

The optical multiplexer performs a function similar to the conventional ignition system: It distributes the output of a pulsed laser among various laser plugs installed in different engine cylinders. The functional requirements of such a high-power optical multiplexer are listed below:

- Ignition timing variation of 0-48° crank angle (CA) before top dead center (BTDC).
- Laser triggers pulse generation.
- Individual cylinder ignition timing variation of ± 6° CA.

Additionally, the system needs to be of low-cost and must conform to the durability targets of the overall system of 80,000 hours.

A survey was performed to evaluate potential technologies that conform to the Aforementioned requirements and the following three were identified:

(i) Electro-optic switches,
(ii) Rotating mirror systems, and
(iii) Flip-flop systems.

The results from the performance evaluation of prototypes of these systems are discussed below. It ought to be noted that the system designs considered below conform to a 6-cylinder, 1800 rpm, 4-stroke engine. These can be extended to engines with more number of cylinders, if necessary.

3.3.1 Electro-Optic Modulator (Pockels Cell)

A Pockels cell acts as an electro-optic switch because it rotates the polarization of the laser light by 90 degrees when activated. As shown in Figure 3.6, a polarizing beam splitter can direct the beam one way or the other based on the polarization state of the laser beam incident on it from the Pockels cell. A linear array of such Pockels cell-beam splitter combinations can be used for the present multiplexing purpose. it demonstrated the required performance (see Figure 3.7).

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Figure 3.6. Schematic of a Pockels Cell-Based Multiplexer.

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Figure 3.7. Photograph of a Pockels Cell-Based Two-Channel Multiplexer.

This scheme facilitates manipulation of the laser beam without having any mechanical movement of the components. However, the number of required optical surfaces increases with the number of engine cylinders and eventually becomes too unwieldy. Additionally voltages as high as 5 kV are required for activating the Pockels cell

3.3.2 Rotating Mirror

This scheme is based on the traditional mechanical distributor system of the 1980’s. As shown in Figure 3.8, a mirror, placed at 45 degrees to the incoming laser beam, is rotated synchronously with the camshaft. As a result, the beam is rotated at 90 degrees about an axis coincident with the incoming laser beam. Laser firing at the appropriate crank angle is facilitated by a timing disk fitted on the shaft. A differential gear system allows ignition timing variation by introducing a phase shift between the mirror and the cam shaft. All of the components required for this system are simple, cost-effective, and offer the required durability. Therefore, a prototype system was designed and assembled. A photograph of the system is shown in Figure 3.9. Tests performed in the lab using a small electrical motor for driving the system showed that it has the required performance. Through such trials it also became evident that significant performance advantages can be gained by using a mirror inclined at angles < 45° to the incoming laser beam, or alternately, by using a small angle transmissive prism. When details of this configuration were presented to engine manufacturers, they stressed the requirement of being able to vary the ignition timing of individual cylinders in natural gas engines. The rotating mirror system does not allow this flexibility. Therefore, development of the flip-flop multiplexing scheme was pursued.

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Figure 3.8. Schematic of a Rotating Mirror Multiplexer.

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Figure 3.9. Photograph of Argonne’s Rotating Mirror Multiplexer.

3.3.3 Flip-flop

A flip-flop multiplexer consists of a linear array of mirrors. An individual mirror, when activated, moves into the path of the laser beam and deflects it into the respective fiber injection port. This system, schematically shown in Figure 3.10, facilitates timing variation of individual cylinders. To test the performance of the flip-flop system, a rotary actuator (Model 6EM) was obtained from the Ledex Division of Saia-Burgess, an automotive parts supplier. This actuator, when activated, rotates its shaft by 22.5° and moves the mirror into or out of the path of the beam. In a 4-stroke, 1800 rpm, 6-cylinder engine the target response time for such a system is 11 ms. The time response of a one-channel system was measured using the arrangement shown in Figure 3.11(a). As noticed in Figure 3.11(b), the system shows very little bounce and in addition, exhibits a time response of less than 7 ms. subsequently, efforts were undertaken to develop a six-channel system for use with a 6-cylinder engine.

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Figure 3.10. Schematic of a Flip-Flop Multiplexer.

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Figure 3.11. (a): Schematic of Setup to Measure the Time Response of the Flip-Flop (b): A Typical Detector Response Curve.

3.4 Fiber-Optic Delivery

As part of the current LIS scheme, it was envisioned that the pulsed output of a laser will be transmitted via optical fibers to laser plugs installed in individual cylinders. The functional requirements specified by engine manufacturers for the optical fibers to be used in a multi-cylinder engine are given below:

- Length > 3 m.

- At least one 90o bend with a radius of curvature < 9.

- Operating temperature > 130° C.

Additionally, to achieve sparking at the distal end of the fiber, a laser flux density of 1012 W/cm2 is necessary when refocused.

For a given fiber and a two-lens system shown in Figure 3.12, the Lagrange invariant yields

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where,

a = core diameter of the fiber (microns)

s = focal spot size (microns)

NAe = numerical aperture at the fiber exit ≈ half cone angle at fiber exit

NAs = numerical aperture at the focal spot

Ilaser = intensity of laser (Watts).

Combining Equations 3.2 and 3.3, the following equation is obtained:

Laser flux density at the focal spot ≈

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In Equation 3.4, the first term is determined by the fiber core material damage threshold. I n the second term, NAs is determined by the space available for the laser plug. For fibers of short lengths (~ 3 m) the numerical aperture at injection end is approximately equal to that at the fiber exit, i.e., NAiNAe; in other words, the exit cone angle is determined by the injection scheme used. Hence the laser flux density can be maximized by making a judicious choice of both the optical fiber and the injection scheme.

However, in deriving Equation 3.4 it has been assumed that the fiber is mode-preserving. For the present case where transmission of high laser power is required, the use of multimode fibers is required. As is usually the case, with the introduction of higher order modes the energy is not equally distributed and is skewed towards the higher order modes. These aspects of actual physical behavior are not captured by Equation 3.4. Nevertheless, this equation provides an acceptable basis that assists in the choice of the optical fiber and the injection scheme.

With the insight provided in the aforementioned arguments, different types of fibers were obtained and tested for sparking in the laboratory. The results of such efforts are discussed below.

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Figure 3.12. Schematic of Laser Refocusing Scheme at the Distal End of the Optical Fiber

3.4.1 Solid Core Fibers

Largely driven by advances in the telecom industry, there have been rapid developments in solid core fiber technology over the past two decades. Currently, step-index (see Figure 3.14(a)) solid core fibers with 1 mm core diameter and a damage threshold of 5 GW/cm2 are readily available. Initial efforts to inject such fibers limited the maximum laser energy transmission to 8 mJ per pulse. The primary mode of failure in such fibers was material damage at the entrance end of the fiber. Analysis showed that significant gains in terms of transmitted laser energy could be obtained by ensuring that the laser beam cross-section profile is incident on the face of the fiber as uniformly as possible. For instance, as shown in Figure 3.13(a), the beam cross-section incident on the fiber face with the use of a simple planoconvex lens had a Peak-to-Average (P/A) ratio of 15.6. However, a profile that distributed the energy more evenly on the fiber face was achieved using an appropriate axicon-lens combination (see Figure 3.13(b)). The P/A ratio in such a case was 3.9 and transmitted laser energy was as high as 30mJ/pulse.

Nevertheless, such advanced injection schemes introduce higher-order modes and the quality of the beam exiting the fiber degrades. Despite the best efforts to counter these effects, the trade-off between the maximum transmittable laser energy and the generation of higher-order modes proved difficult to overcome and sparking could not be achieved. In this respect, the gradient-index fiber shown in Figure 3.14(b) appeared promising as it had better capability to preserve beam quality. Initial trials in the lab showed that these fibers suffer from low material damage threshold. Ultimately, it was decided to direct the present effort toward strategies with hollow-core fibers.

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Figure 3.13. Fiber Face Laser Intensity Distribution Profiles for an Injection Scheme Using (a): Plano-Convex Lens, (b): Combination of Axicon and Plano-Convex Lens.

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Figure 3.14. The Refractive Index Distribution in Two Solid Core Fibers: (a) Step-Index Fiber, and (b) Gradient-Index Fiber.

3.4.2 Advanced Air-Core Fibers

For future on-engine laser ignition applications, two kinds of advanced fiber technologies appear promising: (i) multi-layer hollow glass waveguides, and (ii) hollow-core photonic band-gap fibers. These are schematically shown in Figures 3.15(a) and 3.15(b). Both rely on multiple reflections to confine light to the central air-core, and most importantly both fibers are bend-insensitive. Manufacturing processes of both kinds of fibers are very involved and those that are commercially available – with core diameters <10 μm – are not suitable for the present purpose. It is expected that fibers with core diameters ~ 30 μm will be required to achieve sparking. However, fibers of this size are not readily available.

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Figure 3.15. (a) Schematic of the Cross-Section of a Multi-Layer Hollow Glass Waveguide, and (b) Photograph of an Air-Core Photonic Bandgap Fiber

3.5 Electronic Interface

In natural gas-fueled stationary engines, the Electronic Control Unit (ECU) performs two primary functions: (i) providing ignition with feedback from a timing disk, and (ii) speed control with feedback from a governor. In most of the gas-fueled engines, these functions are performed by two separate units. Earlier in the report, as shown in Figure 3.16, it was envisioned that an electronic interface would be required for the LIS to communicate with the ignition control part of ECU for ignition timing coordination. In consultation with Argonne’s industrial partner, Altronic, Inc., an electronic interface was developed that utilizes commercially available conventional ignition systems.

One such system to drive a flip-flop multiplexer and a laser is schematically shown in Figure 3.16. The position of a timing disk mounted on the crank shaft is sensed by a magnetic pickup. The signal from the magnetic pickup is processed by the ECU/ Conventional ignition system to provide a 165 VDC pulse that is usually fed to the primary side of the ignition coil. In the present electronic interface such a signal is used to generate a 24 VDC pulse using a resistor step down circuit (represented as a χ box in Figure 3.16). The 24 VDC signal, in turn, is used to drive a corresponding rotary actuator. The end of stroke of the rotary actuator is sensed by an optical sensor and a trigger signal is provided to the laser after routing it through an OR gate. In such a system, the individual cylinder ignition timing variation is provided by the ECU/ conventional ignition system. One such system was successfully simulated and tested at Argonne using Altronic, Inc.’s CD-200 system. However the basic design of the system is not specific to a particular model or manufacturer.

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Figure 3.16. Schematic Diagram of the Electronic Interface

3.6 Results and Conclusions,

From an initial survey of several possible LIS schemes, two promising configurations were identified: (i) the laser-per-cylinder concept, and (ii) the multiplexed laser concept, wherein the output of a single laser is distributed over various cylinders. The latter concept was chosen as it promised low-cost and simplicity of thermal management. However, this concept required the development of three high-power components, namely, laser plugs, multiplexers and fiber-optic beam delivery. Guidance was derived in the development of such components through (i) data from fundamental studies conducted in an earlier, and (ii) requirements of the advanced ignition system as specified by engine manufacturers. Specific details of the high-power components considered in the present task are described below:

Laser Plugs: A two-lens design that successfully meets the physical and functional requirements of a laser plug was easily achieved. Adaptation of this design for various engine geometries can be easily achieved.

Multiplexers: Three schemes that distribute the output of a single laser among various cylinders were pursued: (i) an electro-optic switch, (ii) a rotating-mirror scheme, and (iii) a flip-flop switch. The first two schemes fell short of the requirements either due to high-cost or the inability to provide ignition timing variations in individual cylinders. The flip-flop scheme, however, proved effective in all respects.

High-Power Fiber Optic Beam Delivery: Through tests and analyses it was determined that the fiber optic delivery requirements are (i) low divergence at distal end, (ii) high-power laser transmission, and (iii) preservation of mode quality. Initial tests performed using solid core fibers showed that they are limited by the material damage threshold. Subsequent tests performed using Hollow Glass Waveguides showed that they are limited by mode shifts introduced by bending of the optical fibers. While photonic band-gap fibers appear promising, they are not readily available for tests and their development is expected to be expensive.

Electronic Interface: An electronic interface is required for the ALIS to communicate with the ECU of an engine for ignition timing coordination. The timing modules from existing ignition systems were adapted for the present purpose easily.

Chapter 4 Engine Testing Experimental Setup and Procedure

This section describes the engine ignition testing where the test laser has been compared to a commercially available laser. The engine test cell and equipment used to control, monitor, and report the ignition testing are described. The recorded parameters are listed and explained as well as the testing procedure and required modifications to the ignition system.

4.1 Engine Test Cell

Earlier tests performed in an RCM, showed that laser ignition extends the lean ignition limit all the way to the lean flammability limit (ϕ = 0.5) for natural gas-air mixtures. Additionally, it was found that for lean-operation laser ignition accelerates the rate of combustion. The objective of the present task is to determine the impact of such altered combustion behavior on the trade-off between brake thermal efficiency and NOx emissions in a natural gas-fueled engine.

The cell includes fuel flow monitors for both liquid and gaseous fuels. The cell is equipped with temperature controls for the engine coolant, oil, intake air, and the test cell. These controls are vital to the consistency of any tests that are spread over the course of days, weeks or months because they allow for the most similar test situations independent of daily or operational differences. Figure 4.1 shows a basic schematic of the engine test cell.

4.2 Experimental Setup

4.2.1 Single-Cylinder Engine

A large-bore single-cylinder engine was chosen as the experimental engine for the present ignition tests. The specifications of this engine are provided in the Table 3. The engine was coupled to an eddy-current dynamometer. The dynamometer in turn was coupled to a synchronous motor that assisted engine startup. The control scheme and the data acquisition scheme are schematically shown in Figure 4.1

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Figure 4.1: Engine test facility schematic

In-cylinder pressure measurement was performed using a Kistler pressure transducer and combustion heat release analysis was performed subsequently. Emissions measurements were performed using a Horiba emissions bench. Natural gas was injected into the intake manifold using an array of low-pressure injectors and the air/fuel ratio was monitored using a Universal Exhaust Gas Oxygen (UEGO) sensor located on the exhaust manifold.

Table 3. Specifications of Engine

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4.2.2 Open-Path Laser Ignition Setup

For the open-path laser ignition tests, a compact frequency-doubled Nd:YAG laser was directly mounted over the cylinder head. The output of the laser traveled through the valve cover into a laser plug, from where it was focused into the combustion chamber by a sapphire lens. The sapphire lens was mounted flush with the fire deck and was subjected to combustion pressure and temperature. With this lens, the point of laser ignition was centrally located in the combustion chamber at about 13 mm from the lens face, and about 30 mm from the piston at top dead center. The piston, cylinder head, and laser plug are identified in Figure 4.2.

Initial experiments showed that the engine cylinder head vibrated sideways at 15 Hz and ±3 mm amplitude. To minimize the undesirable effects due to these vibrations, the laser was mounted on the ground. The Figure 4.3 shows the mounting scheme and the spark location.

For these tests, Big Sky’s frequency-doubled ‘Ultra’ Nd:YAG laser was used. The Big Sky Ultra laser was capable of delivering 33.5 mJ of energy per pulse at 532 nm when operating at 10Hz. The pulse duration for the Ultra laser was 8 ns. For the present engine tests, the laser output was adjusted to an energy setting of about 29 mJ per pulse. It should be noted that for an equivalence ratio (EQR; previously referred as ϕ) of 0.6 and BMEP of 15 bar, a laser output of about 9 mJ per pulse is sufficient to ensure consistent ignition of the natural gas-air mixture at normal engine operating conditions. The present application required an operating frequency of 3.75 Hz during engine start-up and the ability to transition to a frequency of 7.5 Hz at rated speed (900 rpm). The external trigger of the Ultra laser was utilized in conjunction with the RPECS control system to synchronize the laser pulse to the engine cycle in an appropriate manner. Synchronization also provided the flexibility to adjust laser ignition timing during engine operation.

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Figure 4.2. Photograph Showing the Installed Laser Plug in the Combustion Chamber

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Figure 4.3. Setup for the Open-Path Laser Ignition Tests on a Large-Bore, Single-Cylinder Bombardier Engine.

4.2.3 Fiber-Coupled Laser Ignition Setup

A fiber-coupled laser ignition system, shown in Figure 4.4, was developed for the present subtask. A compact laser that has high beam quality and maximum pulse energy of 90 mJ was used for this purpose. The output of this laser was injected into a 28 inch long, 700 μm core silver/ polymer (Ag/COP) hollow glass waveguide (HGW) using a 250 mm focal length lens.

At the distal end, the laser emission was collimated and refocused using a combination of lenses with focal lengths of 150 mm and 13 mm. To prevent sparking at the fiber input a nitrogen purge was used. A charge-coupled device (CCD) mounted in the proximity allowed fine mechanical adjustments. All of these elements were mounted on a single 1/2”thick plate for testing on the Bombardier CNG engine. Prior to the actual engine tests, engine vibration was simulated in the lab using a shaker table and laser ignition was observed. The fiber-coupled laser ignition system exhibited consistent sparking for over 24 hours. Subsequently, the system was mounted on the engine at 15 degrees from the vertical, as shown in Figure 4.5. A small port on the side of the cylinder head provided optical access for observing the spark kernel.

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Figure 4.4. Layout of the Fiber-Coupled Laser Ignition System

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Figure 4.5. Fiber-Coupled Laser Ignition System as Mounted on the Bombardier CNG Engine.

4.3 Combustion Quality Measurement

For decades researchers have been using in-cylinder pressure information to

investigate the combustion processes in internal combustion engines. The in-cylinder pressure data provides an effective and easy way to monitor and analyze the combustion process on a cycle-to-cycle basis. Current piezo-electric transducer technology offers fast response sensors that easily withstand the harsh environment within the combustion chamber of an internal combustion engine. Voltage signals from the piezo-electric transducer and associated charge amplifier can be sampled and recorded by high speed digital data acquisition systems that are triggered by a crankshaft mounted encoder.

The development of the cylinder pressure in an internal combustion engine is a complex process that involves the heat release from the combustion, heat transfer due to convection, conduction, and radiation, as well as the changing combustion chamber volume. Parameters such as the peak or maximum combustion pressure, (Pmax ), the crank angle where the maximum pressure occurs, (θPmax) and the indicated mean effective pressure (IMEP) can be directly inferred from the in-cylinder pressure data. These combustion parameters are based on the combination of the cylinder pressure data with thermodynamic models. Quantitative measures of the combustion quality can be developed from the use of the in-cylinder pressure data with the cylinder volume and the derivatives of these thermodynamic models. The combustion data that can be extracted from the measured parameters include the heat release rate, ignition delay, and the combustion duration.

The mean effective pressure (MEP) is an important measure of engine performance, because it is relatively invariant with respect to engine size, and it is an excellent measure of the overall performance of the engine. By definition the mean effective pressure is the work produced per unit displaced engine combustion cylinder and is shown in Equation 4.1

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where Wc is the work per cycle and Vd is the displaced volume of the engine. The IMEP is derived from the in-cylinder pressure data and is defined as the work delivered to the piston by the combustion gases during a given engine cycle per unit displaced volume. This quantity neglects all losses (frictional, heat losses, etc.) and measures only the PdV work delivered to the piston. Pumping losses during the intake and exhaust strokes are neglected. The net IMEP is defined in Equation 4.2:

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where P is the in-cylinder pressure (gage) and dVcyl is the differential change in volume. Because the cylinder pressure data is discretely sampled, the integral given in Equation 4.2 can be simplified into a summation of the pressure data collected throughout the compression and expansion strokes. This is possible because the in-cylinder pressure data is discreetly sampled as a function of the crank angle. This transformation is shown in Equation 4.3.

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The resolution of the encoder used to trigger the data acquisition system has been used as the differential change in crank angle, dθ. The volumetric rate of change of the combustion chamber can be derived as a function of engine geometry. The cylinder volume as a function of the crank angle is given by Equation 4.4:

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where Vcyl(θ) is the cylinder volume at any crank angle θ, Vc is the clearance volume of the combustion chamber, rc is the compression ratio, and Rr is the ratio of connecting rod length to crank radius. The cylinder volume rate of change as a function of crank angle is found by differentiating Equation 4.4 with respect to crank angle.

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IMEP as a function of the sampled in-cylinder pressure shown in Equation 4.6

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The net indicated mean effective pressure can be used as a measure of the efficiency of the combustion process and can be easily calculated for any single engine cycle (assuming adequate CA resolution).

The IMEP calculation takes the magnitude of the in-cylinder pressure into account as well as the accuracy of the phasing of the cycle with respect to the maximum brake torque timing. MEP values are frequently estimated from dynamometer measurements of the brake mean effective pressure (BMEP). The BMEP is derived from dynamometer measurements during engine operation and takes into account all mechanical losses. The BMEP of a 4- stroke engine is measured in bar and is shown in Equation 4.7.

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The combustion quality can be represented by the maximum cylinder pressure, which is a very simple means of quantifying the combustion process because it requires only measurements and no calculations. The maximum cylinder pressure for any given cycle is a function of both the total amount of energy released during combustion, and the timing of the energy release with respect to TDC. With all other parameters held constant, the maximum pressure increases as the heat release increases and as the ignition timing is advanced. As the ignition timing is advanced, the combustion heat is released earlier in the cycle, when the cylinder volume is smaller. This leads to higher maximum in-cylinder pressures.

The direct pressure measurements discussed in this section offer a simple method of quantifying the quality of combustion in an engine, however the measurements are affected by several competing phenomenon that are not always easily decoupled. For example, the IMEP describes the total heat release, the timing of the heat release, and the variation in the cycle-to-cycle heat transfer. More specific information is required concerning the parameters that compose the cylinder pressure data. These methods are based on simple combustion models and require simplifying assumptions in order to allow calculation from only the cylinder pressure data. The two most common combustion parameters that are calculated are the heat release rate and the combustion or burn duration.

The rate of heat release is an excellent means by which to characterize the combustion process. Cylinder pressure data can be used to approximate the heat release rate by treating the combustion chamber as a closed system of varying volume during combustion. By neglecting heat transfer and the effects of the crevice volumes, an energy balance on the combustion chamber volume yields the expression in Equation 4.8,

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where dQchem is the incremental amount of chemical energy released, mt is the total mass of the system, Cv is the specific heat at constant volume of the gas, dT is the differential change in the gas temperature, P is the cylinder pressure, and dVcyl is the differential change in the cylinder volume. Equation 4.8 neglects heat transfer, which accounts for approximately 15% of the total heat release. Applying the ideal gas law to Equation 4.8 gives the expression shown in Equation 4.9,

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where Rave is the average gas constant for the cylinder constituents and Cp is the specific heat at constant pressure. Substituting the ratio of specific heats and taking the rate of change with respect to crank angle yields an expression for the heat release rate as a function of the measured pressure, volume, and the rate of change of these properties. This expression is shown in Equation 4.10

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Where γSH is the ratio of specific heats. The ratio of specific heats can be used as a constant or it can be estimated as a function of temperature throughout the combustion process.

The heat release rate can also be integrated and normalized to yield a measure of the burned gas fraction. The importance of the heat release rate is the direct link between the measured cylinder pressure data and the actual cycle resolved combustion events. With correction for heat transfer to the cylinder walls Equation 4.10 can be expressed as Equation 4.11.

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The number of crank angle degrees required for the combustion of the cylinder charge is an important parameter for the operational analysis of internal combustion engines.

The burn duration can be estimated using the heat release rate. The Rassweiler and Withrow mass fraction burned (MFB) model was found to produce the best results in comparative tests (spark ignited engines) with simulated and experimental pressure data. The Rassweiler and Withrow MFB model is shown in Equation 4.12

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Where, MFBθ is the mass fraction burned at crank angle θ, ΔPc is the corrected pressure rise due to combustion, i is the integer crank angle location, ign is the ignition crank angle location and EEOC is the crank angle for estimated end of combustion. The corrected pressure rise due to combustion is calculated from the difference between the incremental measured pressure rise and the pressure rise corresponding to a polytropic compression/expansion process and then referenced to the cylinder volume at TDC and is shown in Equation 4.13:

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Where n is the assumed polytropic index, Vcyl is the cylinder volume and Vr is the reference volume at TDC. Here the polytropic index is based on values calculated from least squares fits to the log pressure versus log volume data usually over a 40 degree period. The two most commonly used burn durations are the number of crank angle degrees between ignition and 2% mass fraction burned, θ 0-2 and the duration of the bulk of combustion, between 10% and 90%, θ10-9.

4.4 Exhaust Gas Analysis

A gas analysis system was used to measure the concentration of gaseous components in the exhaust gas stream. Total hydrocarbons (THC), NOx, CO2 and CO concentrations were measured in the raw exhaust. Three heated stainless steel probes were inserted into the raw engine exhaust to a depth of 1.5" (38 mm) in a 4" (101 mm) exhaust line.

These probes were connected to heated lines which transfer the sampled exhaust to the gas analyzers. The hydrocarbon line and probe were kept at a wall temperature of 375°F±10°F (191°C ± 5.5°C) while the other probes and lines were heated to 175°F±10°F (79°C ± 5.5°C). The temperatures were kept high to prevent condensation of water or volitile organic compounds in the sample lines for the NOx and CO samples and hydrocarbon compounds for the THC sample.

4.4.1 NOx Analysis

Prior to entering the NOx analyzer, significant sample conditioning is required to obtain an accurate and repeatable NOx measurement. The sample line prior to the sample conditioner heated to above the dew point of the sample gas. After traveling the length of the sample line the sample gases first pass through a Horiba Model COM11, NO2 to NO convertor. It is important to convert the NO2 to NO before the water condenser due to substantial NO2 absorption by water. After exiting the NO2 to NO convertor, the sample travels a very short distance to the water condenser. The water condenser used in this study was a Baldwin Model 8210 sample conditioner designed to remove water from the sample stream. This unit is rated to remove less than 10% of any remaining NO2 from the sample stream while removing greater than 98% of the moisture during the condensation process. After moisture removal, the sample enters the NOx analyzer. The indicated concentration was determined to be inversely proportional to the reaction chamber pressure.

It is therefore important to monitor and control the reaction chamber pressure. This is accomplished by using a needle valve at the outlet of the analyzer upstream of the vacuum pump in this study. The ThermoEnvironmental Model 42C operation is based upon the concept that nitric oxide (NO) and ozone (O3) react to produce a characteristic luminescence with intensity linearly proportional to the NO concentration. Infrared light emission results when electronically excited NO2 molecules decay to lower energy states.

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In Equation 4.14, ozone reacts with nitric oxide to produce electronically excited nitrogen dioxide. This NO2* can reach equilibrium either through photoemission (Equation 4.15) or by collisional energy transfer (Equation 4.16), where M represents a neutral molecular constituent within the sample gas mixture. The intensity of the photoemission is given by Equation 4.17:

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where k3M is a function of a specific third body or carrier gas. Because chemiluminescence analyzers measure this intensity, it is advantageous to decrease the probability of collisional de-excitation. This is conventionally accomplished by operation at sub-atmospheric pressures, typically less than 10 mm of Hg. Since k3 is a function of the third body interaction, M, the chemical composition of the carrier gas also affects the measured intensity. NO2 is converted to NO by a stainless steel NO2 to NO converter heated to about 635°C. The sample is then drawn into the analyzer through a particulate filter, a capillary, and then to the mode selection valve. The valve routes the sample either directly to the reaction chamber when the unit is operating in the NO mode or to the NO2 to NO converter and then to the reaction chamber when the analyzer is in the NOx mode. The NOx mode was used in this study. Dry filtered air enters the analyzer through a flow sensor and then through a silent discharge ozonator which generates ozone for the chemiluminescent reaction.

A cooled photomultiplier tube (PMT) detects the NO2 chemiluminescence. The detector operates in pre-set ranges from 10 ppm up to 5,000 ppm. Response time (T90) is 5 seconds. Repeatability and linearity are both less than 1% full scale. Zero drift is less than 0.05 ppm in 24 hours. Span drift is less than 1% full scale per 24 hours. The NOx concentration is transferred to the data acquisition system via 4 - 20 mA analog output.

4.4.2 CO and CO2 analysis

After traveling through a heated filter and heated sample line, sample gases enter a Baldwin Environmental Model 20410 thermoelectric sample gas conditioner which lowers the sample dew point to +5°C (41°F). Particulate matter that may have made it through the upstream heated filter was removed by a Baldwin sample prefilter located downstream from the cooler. The sample once conditioned, is directed to the gas analyzers. The reported soluble gas removal rates as percent of inlet gas concentration are NO: 0% loss, NO2: <2% loss, CO: 0% loss, and CO2: <1% loss. There were two analyzers used in this study, one for CO and one for CO2 in the analyzer bay.

The CO and CO2 analyzers are Horiba Model VIA-510 infrared units. They are based on infrared absorption of the measured component. By determining the amount of attenuation of an infrared beam, the component concentration can be determined. The infrared radiation from a light source is passed through a rotating chopper and into the detection cell. Infrared radiation is absorbed as it passes through the sample. The light then reaches a detection cell which generates an electric current proportional to the incident light. The current is amplified and provided as a 4 - 20 mA analog output. The Horiba VIA-510 has a reproducibility of ±0.5% of fullscale and a zero and span drift of ±1.0% of full-scale per 24 hours.

4.4.3 Total Hydrocarbon Analysis

The sample gases were transferred directly into the heated Horiba FIA-236 flame ionization analyzer. No other conditioning was required because cooling the gas would result in condensation and loss in sample. The analyzer employed a burner/sensor where a regulated flow of sample gas entered a flame produced by reactions of a 40% hydrogen in helium fuel and hydrocarbon free air. The hydrocarbon components of the sample stream underwent a complex ionization that produced electrons and positive ions as shown in Equations 4.18-4.20.

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Polarized electrodes then collect the ions that are produced, causing current to flow through the associated electronic measuring circuitry. The current flow is proportional to the carbon atom concentration. The model FIA-236 measured in set ranges from 10 ppm up to 30,000 ppm, with a T90 of 1.5 seconds. Sensitivity was 0.5% of full scale on each range while repeatability and linearity were both less than 1% full scale. Zero span drift was less than 1% full scale per eight hours. The analog output was 0 -5 VDC non-isolated.

4.5 Results and Discussion

Common metrics for lean-burn engine performance include brake thermal efficiency (η) and brake specific NOX emissions (BSNOx). Macro-operational settings that affect these metrics, ceteris paribus, include fuel-air equivalence ratio (EQR) and ignition timing. The EQR, also written as ϕ, was controlled globally through a fuel metering block and UEGO sensor feedback, but the EQR values reported here were calculated with a carbon-balance method based on exhaust gas analyses [2]. Crank-resolved engine parameters are plotted versus the crank angle (CA) after top dead center (ATDC) at which 50 percent of the total in-cylinder mass had burned (MFB50). This manner of representing the crank-resolved parameters minimized the ambiguity surrounding actual ignition timing in the two ignition systems.

Through combustion analysis of average cylinder pressure (obtained by ensembleaveraging over 100 successive engine cycles), fundamental differences between the two modes of ignition were also compared. Flame development and propagation were quantified through a heat release (burn-rate) analysis. Specific parameters of importance include the crank angle duration for 10 percent, 10 to 50 percent, and 50 to 90 percent of the trapped mass to burn. These parameters are abbreviated as MFB 0-10, MFB 10-50, and MFB 50-90, respectively. Estimating these parameters required an accurate determination of the start of combustion (SOC). For the test results discussed here, SOC was defined as the CA where the heat release becomes non-zero. This method of determining SOC was employed here because in the present laser ignition system the delay between trigger and laser pulse was appreciable (225 μs ± 100 μs). These parameters are delineated in Figure 4.6.

Heat Release / Cylinder Pressure

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Figure 4.6. Arbitrary Cylinder Pressure and Heat Release Comparison to Clarify Nomenclature of Combustion Parameters.

4.5.1 Full Load Comparison (15 bar BMEP)

The engine was operated at a BMEP of 15 bar (or 1307 N-m of torque) at 900 rpm, with both a conventional spark ignition system and an open-path laser ignition system for identifying the benefits laser ignition could have on lean-burn natural gas engines. Laser ignition was found to extend the lean misfire limit from an EQR of 0.55 with open-chamber spark ignition to 0.50. Laser ignition also improved combustion stability at retarded ignition timings. The combined effect of these two improvements was a 50 percent decrease in NOx emissions while maintaining efficiency constant. Figure 4.7 depicts how laser ignition was capable of extending the lean misfire limit. The various data points represent COV of IMEP over a range of EQR for both laser and spark ignition at different ignition timings. In this plot, increased COV of IMEP correlates to more retarded ignition timings at a given EQR.

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Figure 4.7. COV of IMEP Versus Equivalence Ratio (EQR) at a BMEP of 15 bar.

When the engine was operated at increasingly lean EQR values with open-chamber spark ignition, complete misfire occurred at an EQR of 0.55 and was detected by a large COV of IMEP (~15%). Even when ignition timing was advanced to 40° before top dead center (BTDC), a steady torque could not be maintained. By comparison, complete misfire occurred at an EQR of 0.50 with laser ignition. As ignition timing was retarded at a given EQR, the COV of IMEP increased (shown by multiple points in a data set). For both ignition systems, Figures 4.8 and 4.9 show combustion stability as a function of MFB50 at multiple ignition timings. Evidently, laser ignition improved combustion stability at any given MFB50 for every EQR tested.

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Figure 4.8. Combustion Stability with Conventional Spark Ignition at a BMEP of 15 bar.

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Figure 4.9. Combustion Stability with Laser Ignition at a BMEP of 15 bar.

Other higher-energy ignition systems for natural gas engines are also capable of extending the lean misfire limit. These include pre-combustion chambers and high-energy spark ignition systems. However, they do not necessarily improve the overall performance of the engine. Pre-combustion chamber ignition systems can ignite leaner mixtures in the main chamber, but the combustion in the pre-combustion chamber may contribute significantly to the overall NOx emission. High-energy spark ignition systems often accelerate the deterioration of the spark plug, thereby requiring increased maintenance and/or replacement. In addition, they also constrain the flame kernel to the chamber wall where flame growth is often stunted by heat transfer to the metal surroundings. By comparison, laser ignition extended the lean misfire limit, thus gaining the emission and performance benefits without any undesirable side-effects or constraints.

Lean-burn engine technologies are utilized to improve efficiency and decrease NOx emissions. Figure 4.10 compares the trade-off between brake-specific NOx (BSNOx) emissions and brake thermal efficiency for the two ignition systems. This method of comparing the trade-off between engine performance and BSNOx emissions is consistent with other such comparisons of lean-burn technologies. As the engine ran leaner, the trade-off leaned toward lower BSNOx emissions. Laser ignition allowed the engine to run leaner and extended the trend toward lower BSNOx. With open-chamber spark ignition, a brake thermal efficiency value of 32 percent could be obtained along with 17.7 g/kWh (13.2 g/bhphr) BSNOx emissions when the engine was operated at 90 percent of the misfire limit (typical of lean burn applications). Maintaining the same misfire margin and brake thermal efficiency, the BSNOx emissions could be reduced to 5.4 g/kWh (4 g/bhp-hr) using laser ignition; this amounts to a 69 percent improvement. From an alternate viewpoint, while maintaining the NOx emissions level constant at 10 g/kWh (7.5 g/bhp-hr), efficiency improvements up to 3 percentage points are likely with laser ignition.

Combustion characteristics of the two ignition systems were analyzed to examine the reasons behind the benefits observed with laser ignition. Figure 45 shows cylinder pressure data for 100 consecutive cycles for both open-chamber spark ignition (shown to the left) and laser ignition (shown to the right). Spark ignition exhibited complete lean misfire while laser ignition provided relatively stable combustion even at a leaner EQR.

With laser ignition, the COV of IMEP was 5.2 percent at an EQR of 0.52 compared to a COV of IMEP of 16.5 percent at an EQR of 0.55 for conventional spark ignition. Figure 46 shows the ensemble-averaged (over the 100 consecutive cycles) pressure curves for both ignition systems and their corresponding heat release rate profiles. The initial heat release was accelerated with laser ignition and consequently the total burn duration was decreased. As a result of these effects, combustion was more stable and the engine could operate at a lean EQR without misfiring.

For the same reasons, good brake thermal efficiency values were also sustainable at more retarded ignition timings with laser ignition compared to spark ignition. This ability to employ more retarded ignition timings further reduced NOx emissions. The considerable difference in peak pressures between the pressure traces for laser and spark ignition may be attributed to the higher boost pressure required to achieve the same BMEP value (15 bar) at a leaner equivalence ratio with laser ignition.

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Figure 4.10. BSNOX-Brake Thermal Efficiency Tradeoff at a BMEP of 15 bar.

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Figure 4.11. Cylinder Pressure Comparison .

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Figure 4.12. Cylinder Pressure and Heat Release Comparison.

4.5.2 Part Load Comparison (10 bar BMEP)

The performance and emissions improvements shown at full load were also observed at part load. Also, laser ignition was able to extend the lean misfire limit and improve combustion stability; as shown in Figure 4.13. These test results at a BMEP of 10 bar are comparable in which laser ignition extended the lean limit and decreased ignition delay. Using conventional spark ignition, the engine experienced complete lean misfire (COV of IMEP of 15 percent) at an EQR of 0.55. Laser ignition improved combustion stability at the EQR of 0.55 (COV of IMEP of 2.5 percent) and allowed the engine to operate down to an EQR of 0.50, beyond which misfire was observed. Figure 4.14 shows the BSNOx-brake thermal efficiency trade-off at part load illustrates how laser ignition enabled the engine to operate at lower BSNOx emissions without compromising efficiency. The BSNOx emissions were reduced by 65 percent while maintaining the thermal efficiency constant at 32 percent.

illustration not visible in this excerpt

Figure 4.13. COV of IMEP Versus Equivalence Ratio at a BMEP of 10 bar

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Figure 4.14. BSNOX-Brake Thermal Efficiency Trade-off at a BMEP of 10 bar.

4.5.3 Fiber-Coupled Laser Ignition Results

Using the setup shown in Figures 4.4 and 4.5, 100 percent ignition probability was achieved with laser ignition and with the engine operating at an EQR of 0.6 and a BMEP of 10 bar. A comparison of the burn durations for different modes of ignition and different percentage mass fraction burns is shown in Figure 4.15.

This plot shows that the apparent advantages with the use of fiber-coupled laser ignition are similar to those obtained using free-space laser ignition, that is, accelerated burn rates during all phases of combustion. The increased energy content in the plasma kernel generated by laser ignition results in overly reactive species.

These species result in the development of a flame front in ‘overdrive,’ with the associated flame velocities much higher than those typical of conventional spark ignition. The overdriven flame speeds are advantageous to lean-burn engines as the reduced combustion durations can be used to retard ignition, thereby ensuring improved efficiencies and lower NOx emissions.

illustration not visible in this excerpt

Figure 4.15. A Comparison of Burn Durations for Different Modes of Ignition.

4.6 Conclusions

Experiments were performed on a single-cylinder research engine comparing three different ignition systems including:

i. conventional capacitance discharge spark ignition,
ii. free-space laser ignition, and
iii. fiber-coupled laser ignition.

Several benefits were observed with laser ignition including the following:

- Extension of the lean misfire limit by about 10 percent at BMEPS of 10 and 15 bar,
- Increase of the overall burn rate, and
- Improved combustion stability at all comparable test points.

The improved combustion, combined with optimization of engine intake air pressure and ignition timing, was found to result in,

- A reduction of BSNOx emissions by about 70 percent for a given efficiency, or
- An increase in brake thermal efficiencies up to 3 percentage points, while
- maintaining BSNOx emissions constant.

Considering all of the perceived benefits of laser ignition, the on-going efforts aimed to develop an LIS for use with a multi-cylinder engine.

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Details

Title
Laser ignition system on CNG Engine
Course
M.tech IC Engine
Author
Year
2013
Pages
67
Catalog Number
V279136
ISBN (Book)
9783656720058
File size
3156 KB
Language
English
Tags
laser, engine
Quote paper
Vikas Sharma (Author), 2013, Laser ignition system on CNG Engine, Munich, GRIN Verlag, https://www.grin.com/document/279136

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