Fitting curve parameters.
\r\n\tThis project will include chapters covering the main aspects of angiographic techniques; coronary angiography, fluorescein and microangiography, peripheral angiography, miscellaneous angiography, and new concepts and advances. It will provide an insight into significant updates including hybrid imaging, new devices, contrast medium, and techniques. As the endovascular approaches have evolved over the last several years with the rapid influx of minimally invasive techniques, it is important to point out that there are many aspects which require complex medical workups and substantial preoperative decision algorithms, which have not been covered in the literature yet. The book will be a collection of chapters from world class experts contributing to this new endeavor in medical sciences.
",isbn:"978-1-83881-173-0",printIsbn:"978-1-83880-501-2",pdfIsbn:"978-1-83881-174-7",doi:null,price:0,priceEur:null,priceUsd:null,slug:null,numberOfPages:0,isOpenForSubmission:!1,hash:"d6e5b06750aa89961fd7e81c3740c6bd",bookSignature:"Dr. Patricia Bozzetto Ambrosi",publishedDate:null,coverURL:"https://cdn.intechopen.com/books/images_new/9061.jpg",keywords:"Coronary Angiography, Fluorescein, Microangiography, Neurovascular Angiography, Peripheral Angiography, Leg Claudication, Renal Stenosis, Atherosclerosis, Medicolegal, New Advances, Hybrid Imaging",numberOfDownloads:null,numberOfWosCitations:0,numberOfCrossrefCitations:null,numberOfDimensionsCitations:null,numberOfTotalCitations:null,isAvailableForWebshopOrdering:!0,dateEndFirstStepPublish:"September 5th 2019",dateEndSecondStepPublish:"September 26th 2019",dateEndThirdStepPublish:"November 25th 2019",dateEndFourthStepPublish:"February 13th 2020",dateEndFifthStepPublish:"April 13th 2020",remainingDaysToSecondStep:"2 months",secondStepPassed:!0,currentStepOfPublishingProcess:4,editedByType:null,kuFlag:!1,editors:[{id:"221787",title:"Dr.",name:"Patricia",middleName:null,surname:"Bozzetto Ambrosi",slug:"patricia-bozzetto-ambrosi",fullName:"Patricia Bozzetto Ambrosi",profilePictureURL:"https://mts.intechopen.com/storage/users/221787/images/system/221787.jfif",biography:"Dr. Patricia Ambrosi graduated from the University of Caxias do Sul, Brazil, and University of Rome Tor Vergata, Italy. She is a former Researcher in the area of Morphophysiology at the University of Cordoba and in the area of Pathology at the Reina Sofia Hospital, Córdoba, Spain. She was a medical trainee in Neurosurgery at the University of Tubingen, Germany. She was a specialist in Neurology and Neurosurgery ( Hospital of Restauração), Neuroradiology and Imaging Diagnostics at Paris Marie Curie University. She holds a Master degree in Medicine from the Universidade Nova Lisboa in Portugal and in Behavioral Sciences and Neuropsychiatry at UFPE. She made sandwich Ph.D. in Biological Sciences at UFPE and Paris Diderot University. She was Assistant Physician in Interventional Neuroradiology at the Ophthalmological Foundation Adolphe de Rothschild, at Beaujon Hospital and Assistant Physician in Diagnostic Imaging at Hospices Civil de Strasbourg. She was Praticien Associé in Interventional Neuroradiology at Neurologique Hospital Pierre Wertheimer, University of Lyon Claude Bernard in Lyon, France and She is actually Independent Consultant in Neuroradiology, Endovascular Surgery, and Imaging Diagnostics. 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From chapter submission and review, to approval and revision, copyediting and design, until final publication, I work closely with authors and editors to ensure a simple and easy publishing process. I maintain constant and effective communication with authors, editors and reviewers, which allows for a level of personal support that enables contributors to fully commit and concentrate on the chapters they are writing, editing, or reviewing. I assist authors in the preparation of their full chapter submissions and track important deadlines and ensure they are met. I help to coordinate internal processes such as linguistic review, and monitor the technical aspects of the process. As an ASM I am also involved in the acquisition of editors. 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Chan and Manoj Kumar Tiwari",coverURL:"https://cdn.intechopen.com/books/images_new/3794.jpg",editedByType:"Edited by",editors:[{id:"252210",title:"Dr.",name:"Felix",surname:"Chan",slug:"felix-chan",fullName:"Felix Chan"}],productType:{id:"1",chapterContentType:"chapter",authoredCaption:"Edited by"}}]},chapter:{item:{type:"chapter",id:"60161",title:"HyDRa: Vortex Polishing with a Deterministic Hydrodynamic Radial Polishing Tool",doi:"10.5772/intechopen.75524",slug:"hydra-vortex-polishing-with-a-deterministic-hydrodynamic-radial-polishing-tool",body:'As the demand to fabricate larger and increasingly complex aspheric optics becomes more common, the need for deterministic polishing tools which can meet these demands has motivated the development of the hydrodynamic, polishing tool (HyDRa).
The HyDRa polishing tool [1, 2] is a non-contact, zero-force hydrodynamic tool that rotationally accelerates a slurry and air mixture and expels it tangentially onto the workpiece. It consists of several operational stages: an abrasive suspension (water and polishing grit) is fed to the tool’s first operational stage, where it is mixed with air at a controlled pressure in order to produce a variable-density abrasive foam. This foam then enters the rotational acceleration chamber, where it is sped up to high revolutions per minute (RPM). The rotational energy of the flow then turns into radial velocity in a nozzle that forms between the tool’s divergent output and the workpiece. In this way, the abrasive particles graze the workpiece in closely tangential trajectories. Thus, the polishing particles generate a shearing action that removes material in a ductile removal process, as described in [3]. The drag generated by this radial flow forms a central low-pressure zone (vacuum) surrounded by a high-pressure ring. As a result, the tool floats over the workpiece, exerting no net force onto it [4], since these regions cancel each other out. This poses many advantages in modern deterministic polishing, in particular the capability of polishing ultra-thin surfaces, such as semiconductor wafers and optical membranes.
The HyDRa tool belongs to the fluid jet polishing (FJP) family, originally developed by Fähnle et al. [5]. Although HyDRa also expels an abrasive suspension onto the surface to be polished, there are several basic operational principles that differ from the classic FJP technique. The FJP method, and most other polishing techniques, needs to apply pressure onto the workpiece, in order to remove material. In some cases, as in the classic FJP technique, the force that is exerted onto the material can be minimized by using a small contact area; see for example [6], where the force on the workpiece is less than 1 N. This however, could represent a trade-off between footprint size and removal rate; removal rates of <0.01 mm3/h are common for 1–2 μm Cerium oxide polisher [6]. Other deterministic methods, such as ion beam figuring (IBF), are non-contact, zero-force processes that present removal rates of up to 50 mm3/min [7] without degrading the initial micro-roughness.
This section describes the generalities of polishing with the HyDRa system. A typical static removal footprint is shown in Figure 1.
Tool influence function. The image to the right shows the removal function profile extracted from the interferogram to the left. This profile was obtained by operating the tool for 2 s on a fixed position over the workpiece. DH is the full width at half maximum (FWHM) of the removal function. Tool footprint diameters typically range between 3 and 10 mm, depending on the particular tool.
As can be seen, the removal takes place in a ~5 mm diameter region that has an M-shaped geometry. Removal in the central region is about 20% lower than the peripheral region. Removal is a function of slurry density (ρ) and accelerating air pressure (PT), slurry flow, i.e., flow rate (F1), and is a linear function of dwell-time. In order to attain deterministic polishing, all tool parameters are controlled with high precision, as described in Section 4.
The HyDRa tool is part of a complex polishing robot, consisting of a CNC positioning device; a fluid and compressed air control system; a slurry management unit, which stirs the slurry and controls its density; and a software package that obtains error maps from a series of interferometers, and generates dwell-time/constant-velocity PWM trajectories. These trajectories correct the workpiece figure, depending on the selected method (pulsed or continuous operation), as the block-diagram of Figure 2 illustrates.
HyDRa polishing process setup, (see text).
The HyDRa tool is attached to a five degree of freedom (DOF) polishing machine with force feedback, based on a 2.4 × 2.4 m Cartesian CNC, with two additional DOF (tip and tilt), implemented by means of a 3-actuator hexapod. This configuration allows the generation and polishing of any surface geometry. Since feedback control keeps all polishing parameters constant, removal is exclusively a function of dwell-time. In this way, figure correction depends on the trajectory followed by the tool and the velocity at each point along it, which requires that the CNC be capable of following five-dimensional, controlled-velocity trajectories. Simultaneously, z-axis movement is controlled so that the tool can accurately follow the surface contour with zero-force. Although the machine’s repeatability is around 10 μm, the removal is accomplished with nanometric accuracy due to a load cell that regulates tool height over the workpiece, as reported in [4]. The slurry conditioning unit (SCU) supplies a density controlled polishing suspension to the HyDRa slurry supply system. It also captures and reincorporates the liquid and atomized slurry that is expelled from the tool. The density is continuously monitored by means of a photodensitometer and controlled by means of polishing-paste and water supply systems. The expelled slurry is captured by the return system, which consists of a blower that reincorporates this slurry and air into the SCU container, forcing the mixture through a washed-air system.
A pump is continually recirculating the polisher in the container, and a derivation supplies filtered slurry to the HyDRa slurry control system. The slurry that flows into the tool is fed by means of a damped-diaphragm DC pump that is feedback-controlled by means of either a flow meter or a pressure sensor. In order to reduce air contamination, the HyDRa system polishes in shallow immersion so that an air-lock for the return system is formed. The HyDRa tool accelerates an abrasive foam, composed of a variable-density suspension of slurry and air created by means of an air control system that regulates the foaming and propelling air pressures using electromechanic air regulators, pressure sensors and control electronics. All polishing parameters are acquired with a data acquisition card, and controlled and visualized in LabView®. The surface is fixed onto the CNC platform as the tool is swept over the surface in a pattern chosen by the user. Pulsed operation of the tool sweeps the tool(s) along the workpiece at a constant velocity, switching the polishing action on and off depending on how much material needs to be removed. The slurry is fed to the tool at controlled pressures and flow rates. The tool is supplied with compressed air to operate the foaming and acceleration stages. The abrasive foam is radially expelled through the tool nozzle onto the surface to be polished, as described above. Tool height is controlled by the feedback variable provided by the load cell, so that the CNC can adjust tool height in order to polish with zero force, as is discussed below.
One of the central advantages of the HyDRa tool resides in that it can be adjusted to exert zero force on the work surface, while maintaining considerable removal rates (~10 mm3/h). The flotation effect is described in more depth in this section.
As the HyDRa tool expels the slurry through its nozzle, the rotational energy of the flow is converted to radial velocity. The drag generated by this radial flow produces a central low-pressure zone that is surrounded by a vortex which is in turn confined by a ring shaped, positive thrust zone. This effect differs from the linear jet polishing technique [3, 8] in that, in normal-incidence, classic jets, the pressure profile is represented by a Gaussian distribution, where the maximum is located at the jet’s center. In contrast, the pressure distribution of HyDRa on the workpiece presents negative values at the footprint’s center and is circumscribed by an annular, positive-pressure region. With this, the force on the workpiece can be adjusted to net zero-force. In Figure 3a, footprint pressure vs. distance from the center of the tool is plotted. This was obtained by means of a 0.46 mm diameter orifice connected to a pressure sensor over which the tool footprint was radially scanned. As can be seen, a low pressure zone forms at the central part of the footprint and, as the radius increases, the gauge pressure increases to a maximum. It then falls again as the orifice approaches the tool’s outer radius.
(a) (left) Pressure distribution as a function of tool radius for different operating pressures. (b) (right) Force applied by the HyDRa tool as a function of distance Z from the work surface for accelerating pressures PT = 20, 30, 40 and 50 PSI. For clarity, a constant offset has been added to the 30, 40 and 50 cases. Typical error bars are 0.1 N for the load cell and 20 μm for the Z position. The solid lines are an analytical function fit derived from a Morse potential. In the upper right inset, the restitution spring constant is shown to be related to PT. When an external holding force Fh is applied to the tool and pushes it against the workpiece, the vacuum force Fv has a tendency to decrease as the thrust force Fe is increased.
As the tool moves away from the workpiece, the vacuum has a tendency to increase, as the thrust force tends to decrease; in consequence, the balance of these two forces is a function of the tool’s distance from the surface to be polished. With this, several operational modes can be achieved, one of them allowing the tool to freely float over the surface without the need to rigidly attach it to a positioning device. In this case, the tool’s weight is counterbalanced by the thrust and vacuum forces. When the HyDRa tool is mounted onto a passive hexapod through a load cell, the support force Fh counterbalances the tool weight mg and the remaining forces to zero, while maintaining the tool in a static position.
The load cell value Fh is then used as feedback for the control system to maintain zero force on the workpiece. This is done by adjusting the z-axis of the CNC/hexapod.
We measured the tool and workpiece force interactions as a function of separation. In Figure 3b, the overall behavior of this interaction is shown in experiments where slurry accelerating pressures of 20 to 50 PSI and a polishing slurry flow of 5 ml/s were used. The distance over the workpiece was modified from several millimeters, down to a few hundred microns. The values of the load cell Fh were used to obtain the substrate force Fx, as explained above.
When the tool exceeds a certain distance, the workpiece experiences no force (see Figure 3). As the tool approaches the substrate, a negative force (attraction) is produced by a vacuum that develops between the workpiece and the HyDRa tool. As the tool further approaches the workpiece, a positive repulsion force is experienced, caused by the thrust force. At a few hundred μm from the workpiece, the vacuum and thrust forces balance each other out to zero. Here the surface can be polished without being deformed by the tool.
The combination of attractive and repulsive forces commonly arises in physical problems. One of such is given by the empirical Morse potential of diatomic molecules
50 | 10 | 1.36 | 0.36 | 40 |
40 | 7.5 | 1.36 | 0.34 | 27.7 |
30 | 5.1 | 1.36 | 0.34 | 18.9 |
20 | 3 | 1.36 | 0.34 | 11.1 |
Fitting curve parameters.
Several tools have been developed to accommodate for different removal rates that range from 1 to 600 mm3/h. This section presents measurements of static and volumetric removal rates for a medium removal-rate tool with a nozzle diameter of 3.8 mm. Several polisher grits have been tested: Cerium oxide (Opaline) with a particle size of 1 μm, and aluminum oxide (μ-grit) with particle sizes of 5 and 12 μm, suspended in water at a constant relative density of 1.09. Samples of different materials were polished: standard window glass, water-free fused silica (Infrasil®-302), BK-7 borosilicate, and Ohara’s CLEARCERAM-Z® vitroceramic.
The measurements were performed by scan-sliding the tool over the sample at a constant speed and a corresponding zero-force tool height. The resulting cavity was measured interferometrically, and both the volume of removed material and the depth of the cavity were calculated. With these measurements the static removal rate DS (depth/time) and the volumetric removal rate DV were then calculated. We performed these measurements at different operational conditions of accelerating pressure (PT) and slurry flow (fp). In all cases, the relative density of the slurry was kept at 1.09. In Figure 4, the test results are shown. In all three panels, the left vertical axis DV is related to DS as DV = DS* A, where A is the tool footprint area. Each plot indicates the dependence of removal on the accelerating pressure PT. The top left panel, where Cerium oxide was used, slurry flow fp values of 3, 5 and 7 ml/s on window glass were tested. These are indicated as numbers on the curve.
Volumetric removal rate DV as a function of accelerating pressure (PT) for various polisher types: 1 μm cerium oxide (Opaline), 5 μm aluminum oxide (μgrit) and 12 μm aluminum oxide. Opaline removal rates are presented for window glass (solid line), fused silica, and borosilicate (dotted line). Removal of μ-grit is shown for window glass and CLEARCERAM®-Z vitroceramic. All fitted curves are versions of the same second order polynomial law, scaled by a factor that is a function of substrate material and particle size (see text). Zero-force polishing was achieved at the corresponding tool height for this effect.
As can be noted, the highest removal rate for this tool is achieved with a slurry flow of 5 ml/s, thus this flow was employed for all consecutive experiments. Removal rates up to 2.5 mm3/h at 50 PSI are obtained for this polisher. The fp = 5 points are well fitted by a second order polynomial. In the first panel we also show the results of polishing a fused silica sample (crosses). In this case the hardness of the material causes the removal rate to decrease to 66% compared to window glass. The dotted line that fits these values is obtained by multiplying the same DS polynomial by a small factor. In the lower left panel we tested the removal rate of 5 μm aluminum oxide grit on glass (crosses). The volumetric removal rate rises to ~50 mm3/h at an accelerating pressure of 50 PSI. The solid line is the DS polynomial obtained previously, multiplied by 19. Again, a very good fit can be seen. Lastly, in the lower-right panel, the removal rate of aluminum oxide grit (12 μm particle size) on Ohara CLEARCERAM® vitroceramic is shown. Removal was so large that it was impossible to measure it reliably with an interferometer, so the resulting cavity was measured by means of a needle profilometer. Only two points were taken in this experiment (crosses in the last graph). If the same polynomial DS were to be fitted (dashed line), it would have to be multiplied by 94. For all cases, the removed volume per abrasive particle is 3 to 5 orders of magnitude lower than the particle’s volume. The kinetic energy as well as the grazing incidence associated with both the cerium and alumina particles, creates stresses that are not large enough to produce permanent dents. Resulting typical micro-roughnesses of 2, 24, and 31 nm, for grit sizes of 1, 5, and 12 μm, respectively, are around three orders of magnitude smaller than the particle sizes. These tests demonstrate that independently of grit size and material, as well as substrate hardness, behavior of removal as a function of accelerating tool pressure is comparable, which indicates that the same material removal process discussed above is taking place. The scaling parameter is an indicator of how efficiently a particular abrasive removes material from a surface of a given hardness. All HyDRa tools present this same relationship of removal as a function of PT and grit size for pressures of up to 90 PSI. Currently, typical removal rates of ~15 mm3/h @ 90 PSI, using 1 μm Opaline on borosilicate glass, are common.
The HyDRa trajectory planning tool (HyTPT) is a software package developed specifically for the HyDRa tool [9]. It feeds machining code to a CNC or any computer-controlled positioning device, based on the error map that has been obtained interferometrically. When only one HyDRa tool is available, the amount of material that is removed at each specific position is proportional to the dwell time. This dwell time can be controlled either by the speed of the tool along a given trajectory when the tool is operated in continuous mode, or by the width of the pulses when operated in pulsed mode.
In HyTPT (Figure 5) a main window presents the project name and grants the user the ability to go from error maps to machine coordinates in four steps:
Error Map alignment: This allows the easy alignment of the error map, and the determination of its center, orientation and pixel size. The basic shapes are rectangular, circular, and annular surfaces.
Base trajector: It is possible to select the base trajectory from a series of curves: from simple raster patterns to more complex curves, such as trochoids and rotating triangles. It is possible to define and analyze each curve from the point of view of trajectory density, total polishing time, and speed at each point, among several others.
Surface shape: The shape of the surface, currently any on- or off-axis conic section, is defined in this window. The 2-D base trajectory is projected onto this 3-D surface. In addition to the 3-D trajectory the normal vector to the surface at each position is calculated, so that the CNC may orient the HyDRa tool normal to the surface along the trajectory.
Machine coordinates: A set of alignment tools facilitates the transformation of surface coordinates to machine coordinates. These routines communicate with the CNC machine and are developed for each particular CNC or robot arm to which the HyDRa tool is attached. Finally, the centering and orientation parameters, nine in total, are determined in order to properly align the system. Once each of these parameters has been calculated, the trajectory is delivered to the CNC machine, or robot arm so it can be executed there.
HyTPT software main interface showing all four steps.
Due to its hydrodynamic properties, the original HyDRa tool [1] does not allow the modification or switching of any of its operational parameters during operation. Therefore, in order to control the removal, dwell time is modified by means of tool velocity. However, there is a maximum tool velocity imposed by CNC limitations, that fixes a minimum, non-zero amount of material that can be removed with the tool. This poses problems for several operational applications, such as zonal corrections, multi-head or tessellated polishing, and edge problems. A new HyDRa tool design (patent pending) [2] overcomes these problems, by switching one of the operational parameters (polisher flow). This enables the pulsing of the abrasive action at will, without affecting tool bias. The ability to operate the HyDRa tool in a switched manner widens its overall performance and efficiency, adding new applications. It is not simple to switch most of the tool’s operational parameters during use, since this causes a loss of tool bias, which affects the flotation capability. Restoring these parameters on the fly takes time and produces unwanted effects such as cavitation. This has limited the polishing strategy to making full sweeps of the entire surface. When needing to polish only a small section, the region must be approached with the tool biased, leaving behind an unwanted track and approach and exit marks. A new HyDRa tool design has been developed which allows switching slurry flow without losing tool bias. Switching frequencies of up to 10 Hz, and pulses as narrow as 10 ms, can be achieved with this new tool. This is accomplished by means of an overdriven electro-valve which is installed in, or close to, the tool. This allows the use of polishing pulses that can be applied on a per-pixel basis or in a continuous scan using pulse width modulation (PWM) techniques. With this feature, dwell time can be controlled below the minimum attainable by a continuous action at the maximum CNC speed.
A pulsed HyDRa tool with a 7 mm footprint has been developed and tested for linearity. Figure 6a shows the results of erosion vs. dwell-time. Pulse width was varied at constant increments, starting from 10 ms to a maximum of 500 ms, as the tool was moved at 0.2 mm increments, overlapping 35 times at each tool footprint diameter. The erosion was measured using a Fizeau interferometer and the result was normalized, so that removal corresponded to a single pass of the tool over each point along the line that was polished. Error bars are primarily due to errors produced by the subtraction of the base reference during interferogram reduction. A removal resolution of 0.1 nm/ms can be seen from the data. Noticeable polishing effects were observed at 25 ms. We attribute this effect to the electro-valve response time, which can be improved by using faster actuators.
(a) Erosion vs. pulse-width in a linearity test. (b) Response function of the tool for pulsed operation at constant tool velocity. The tool was switched in a PWM mode, with a 50% duty cycle and 100 ms on/off switching times. The raster pattern used to generate this surface is perpendicular to the surface produced in the figure. Pixel size is limited by footprint size. Tool velocity was chosen so the pulses could be resolved independently.
It is now possible to control the duration of a pulse as a fraction of the time it takes the tool to travel a distance of one footprint diameter.
The depth of removed material for a raster scan pattern h, can be given by:
where DV is the volumetric removal rate, Y is the PWM duty cycle (ON time divided by the period T), V is the CNC velocity, and S is the raster step size. The period of the switching signal is T = D/V, where D is the tool footprint diameter. In the time it takes to cross a tool footprint, there is a single pulse whose width can be varied from zero to the entire footprint diameter. The depth of the removed material h = βτ is proportional to the dwell time τ, defined by τ = YD/V, with a proportionality constant β = DV/SD. When Y = 1, which is the continuous mode, tool removal (dwell time) h is controlled by CNC velocity. When Y < 1, which is the pulsed mode, h is controlled by means of switching the slurry supply and keeping the tool velocity constant. The continuous mode is limited to removals greater than h{min} = DV Y/V{max}S. In order to obtain a lower removal, the pulsed mode must be employed.
For a tool footprint diameter of 7 mm and a maximum CNC velocity of 2000 mm/min, the minimum period of the switching signal will be 0.2 seg, or 5 HZ, which is within the 10 Hz switching frequency range. In an extreme case, where the duty cycle is switched between 0 and 1 while maintaining a constant velocity, it is possible to create a pixelated pattern that is useful for determining the response function of the pulsed tool (Figure 6b). A fringed pattern can be observed where the interface between the regions presents a slope that corresponds to the tool footprint diameter, which is the limiting polishing element size (poxel, or polishing element).
When only a small section of the surface needs to be polished, this region must be approached with the tool turned on, leaving behind unwanted tracks, as well as approach and exit marks. This is solved by using the pulsed mode of the HyDRa tool. When an isolated region that needs further polishing is identified, a dampening band of constant width surrounding it is defined. Assuming a raster pattern is used, the region is approached with the HyDRa tool fully operational with Y = 0, until it enters the dampening region. Here velocity is smoothly incremented to the value needed inside the region while at the same time, the desired dwell time is controlled by means of PWM. The width of the dampening region is determined by the CNC acceleration and deceleration capabilities. Inside the region to be corrected, either a pulsed or continuous polishing can be used in order to maximize efficiency.
In the constant velocity PWM polishing case, the resulting response function in the sweep direction is different from the response function of the transverse, raster direction. When a symmetrical finishing is needed, it is possible to employ the pixel polishing method, which consists of stepping the tool at discrete positions with respect to each other, covering the region of interest with the same step increments in both axes. The tool is then switched on for the necessary time in order to achieve the desired removal for each position. This method can also be useful when very localized zonal polishing is needed. This allows the tool to either follow a raster pattern, or any other trajectory or set of discrete positions over the region of interest. An example of this method was used in Section 3.1 for the linearity test, where it was shown that a removal resolution of 0.1 nm can be attained.
In the polishing of meter-class surfaces, efficiency is limited due to HyDRa’s small footprint size and volumetric removal rate. Efficiency, however, can be improved by simultaneously polishing the surface with several HyDRa tools. These tools can be mounted on independent polishing robots, where each robot tackles a certain section of the surface. Alternatively, several tools can be mounted on a single robot arm, as described below. This method poses several problems, such as obtaining smooth seams between sections, approaching each section without leaving marks, and avoiding collisions as two tools concurrently approach the boundary between sections. In order to obtain a seamless interface between two independent sections, it is necessary to approach the boundary following special trajectories, such as the wedge pattern shown in Figure 7 (bottom). This trajectory avoids duplicating dwell time at the seam, such as would happen if a rectangular pattern were used (top of Figure 7).
Tessellated polishing showing two different trajectory approaches. The figures to the left show the region of the interface between two raster patterns that use a rectangular (upper) and wedge (lower) seam at the boundary. The middle figures show a simulation of the resulting removal with the tool moving at a constant velocity and without PWM. Note the approach/exit tool mark at the beginning/end of the trajectory due to not switching the tool off. The interferograms to the right are actual polishing experiments. As can be seen, the raster pattern using rectangular trajectories duplicates removal at the seam, whereas the wedge trajectory produces a seamless interface. Another possible solution is to vary the pulse width at the seam, in order to match the dwell time between adjacent polishing sections.
When the polishing process needs to be interrupted, it is now possible to stop at any point on the surface and continue polishing at a later time.
As in other polishing methods, the HyDRa tool tends to leave a small (one footprint diameter) fallen edge. In order to overcome this with the unmodified HyDRa tool, tool velocity is incremented as the tool approaches the edge, reducing dwell-time. This is counterintuitive and can present CNC control problems, since the tool is accelerated in a region where it should be preparing for a raster direction change. The on–off capability of the new tool can alleviate this problem, since dwell-time can be controlled without having to increment tool velocity at the edge region. In fact, this method allows for decelerating the CNC in order to prepare for a direction change.
Another advantage of being able to pulse the tool is a quicker convergence towards the desired surface [10]. As pointed out in [11], the existence of a minimum amount that will be removed due to not being able to turn the tool off (hmin), limits the amount of material that can be removed in each run, whereas, by being able to pulse the tool, hmin can be made zero, allowing for a maximum value of f, further increasing the polishing convergence rate.
Since HyDRa tools can now be pulsed, several polishing heads can be mounted onto a common arm which moves at a constant velocity over the surface. Dwell time is then controlled using PWM for each tool, as required by the error map. We can also take advantage of the self-conforming capabilities of HyDRa, in that it is not necessary to employ a positioning device to conform the parallelism of the tool to the surface. Only one degree of freedom (DOF) per tool is required. Each loop is closed with a load cell signal and implemented by means of a linear stage, which permits zero-force polishing while freely following the local sag and tilt of the surface. Another advantage of this type of polishing is that one single slurry supply system can be used for all the tools, simplifying the system and considerably reducing the costs. Polishing efficiency becomes a function of the number of tools, and in the case of a matrix configuration, several polishing runs can be implemented into a single sweep, reducing polishing time.
By polishing with several tools, each tool is essentially given a section of the surface and the boundaries between sections are finished seamlessly, either by employing wedged joints, or by using PWM. Among the possible multi-tool configurations are matrix, linear and spiral layouts:
By mounting several HyDRa tools onto a single polishing arm, attached to a Cartesian CNC machine, it is possible to cover an area by sweeping the arm in the x and y directions. Each tool is separated from the next by a fixed distance δ in the x axis. The sweeping action in the x axis is done by moving the arm by δ, and then advancing with the selected raster step in the y axis. The overlap between the sections assigned to each tool is managed by either using tessellated or the PWM techniques, as described above. There are certain considerations to be taken into account for this method, particularly due to the edge problem that arises when polishing circular or non-rectangular surfaces. There will always be a tool that needs to either enter or exit the surface, while others are already polishing. Additionally, since these tools need to take advantage of their self-conforming capability, as they approach the edge of the workpiece, they loose floatability. These problems can be dealt with by adequate trajectory programming.
The linear configuration can be expanded by creating a matrix of tightly packed HyDRa tools maximizing tool number in order to minimize polishing time. The working principle is the same as the linear case, but adding M rows. This is equivalent to carrying out M polishing runs in a single iteration.
In the case of large circular mirrors, it may be more efficient to polish the mirror by placing it on a rotary table. If we seek to fix as many HyDRa tools as possible onto a single arm, two conditions must be met. First, the number of tools should increase as r2, so that each tool covers the same mirror area. Secondly, tools must be packed at the maximum allowable density, so that each tool sits next to the following one. Therefore, the shape of the arm must be a kind of spiral which is possible to solve for. By calculus of variations a spiral curve parameterized by the (increasing) radial coordinate, given in polar coordinates (r, θ) results in:
Hydra tools mounted on a spiral arm (see text).
Of the multiple finishing techniques currently in use, the ones based on sub-aperture polishing may be candidates for deterministic polishing, provided that the uncertainty of key polishing parameters is minimized. Deterministic polishing relies on a stable and predictable tool influence function; thus it is imperative that it is fully characterized for each material that will be polished. Simultaneously, metrology is a determining factor of the final quality of the surface, since it limits the precision of the error maps that can be obtained. This requires the knowledge of a series of polishing parameters such as tool velocity, pressure and height as well as slurry type, temperature, etc. Most of these parameters remain constant during the time periods required for polishing small optics, i.e. a few minutes. If larger, meter-class surfaces need to be polished, it is important to control and keep all parameters constant during an entire polishing run, which can represent over 10 h. Thus, a very stable and precise process control of the process is required.
The HyDRa tool removal function is based mainly on four independent operating parameters: propelling air pressure, grit mass concentration, height of the tool over the surface to be polished, and slurry flow and/or slurry pressure. In order to ensure deterministically polished surfaces, the errors contributed by each of these factors must be taken into account and precisely controlled to 1% for the entire length of the polishing run (over 100 h). Simultaneously, metrology is crucial for determining the surface’s final quality, since it dictates the limit of the precision of the error maps that can be obtained.
To maximize polishing performance, an abrasive foam is created in the tool’s first stage. This raises the velocity of the polishing particles, improving the removal of material. This foam is produced by combining a constant flow f (a few ml/s) slurry, with air that is kept at a constant pressure Pp. This fluid is then accelerated with pressurized air at a propelling pressure PT, in one or more cylindrical cavities. The resulting abrasive foam is then expelled through the tool’s nozzle, where a vortex is produced that develops into a radial flow, and generates a grazing, uniform removal footprint. A relation of slurry flow f to slurry pressure Pp exists for each value of accelerating pressure PT. This, in addition, depends on the tool’s physical characteristics, such as its overall dimensions, the geometry of the acceleration chamber(s), as well as the nozzle shape. This relation establishes an operational diagram that defines tool bias. In this section, the control of f is chosen, although it is possible to select to control for either f or Pp. The removal D of HyDRa mainly depends on four independent operating parameters: propelling air pressure PT, grit mass concentration ρi, slurry flow f, and distance of the tool over the workpiece Z. In order for deterministically polished surfaces to be obtained, the errors contributed by each of these parameters must be taken into account and controlled.
The removal rates, as determined by a series of independent experiments, where the polishing parameters varied, are shown in Figure 9.
Normalized removal rate D/D¯ (dimensionless) with reference to the normalized value of certain polishing parameters (mass concentration ρi, propelling pressure PT, slurry flow f and tool height z). The sensitivity of removal to each parameter is indicated in the upper left corner of each graph.
To generalize the analysis, all parameters X are normalized around their operational values as
In the case of tool height z, load cell force Fc can be used instead, since, as shown in [4] tool force is an approximate linear function of distance when close to the operation point, given that z + K Fc, with K~10 𝜇m/N, and hence δz = δFc..
If we assume that each of these four variables is statistically independent, the total error can be added in quadrature. For example, if each parameter is controlled to ~1% precision, then the total error δDT is
As mentioned before, in order to deterministically polish large surfaces it is imperative that removal rate remains stable over extended time periods. We polished an 84-cm hyperbolic primary mirror to λ/10 RMS, 0.7 λ PV in order to prove that HyDRa could deterministically tackle meter-class optics. The polishing process is described in [12]. From the error maps that were acquired during the iterations, the level of determinism of the process could be calculated. From each map we computed a tool trajectory with distinct dwell times. The amount of removed material was calculated by subtracting the previous error map from the measured one. Then, from the obtained result after polishing, the removed material for each iteration was determined and plotted as a function of dwell time. Refer to Figure 10. A linear relation is expected and the deviation from this represents the level of determinism, Figure 11. This experiment was useful to evaluate the importance of the stability of each parameter in the level of determinism for prolonged time periods. In the figure, a larger error can be noticed for shorter dwell-times than for longer ones. This is due to CNC errors when the tool has to be quickly accelerated to obtain short dwell-times. As the mirror is progressively corrected, the surface is smoother and these changes tend to decrease.
An 84 cm mirror with a 1 cm thick faceplate that was polished using the HyDRa system. (a) Picture of the mirror’s internal back-structure. (b) Mirror prior to HyDRa finishing. The print-through left by the original lap-polishing process can be noted. (c) Mirror surface after HyDRa polishing. The polishing process entirely removed the print-through by polishing with the zero-force, error-map based process described in this chapter. Low-order Zernikes have been removed so this effect is highlighted. Z-scales are the same and are shown as vertical bars in nm.
Total amount of material removed in the final iteration (3 runs = 30 h), as a function of dwell time in each area element of 2.6 × 2.6 mm2 (pixel size). A linear relation at a removal rate of 13 mm3/h (shown as a solid line) is expected for an entirely deterministic process. The true deviations from this behavior amount to 10.6%, which is the attained level of non-determinism and represents the standard deviation of the points with respect to the best-fit line.
Three components in the power spectral density (PSD) of the residual surface errors that are related to the footprint diameter of the tool DH exist for any given polishing method. In the low frequency domain (L >> DH), the surface errors (optical figure) are a function of the stability of the polishing parameters during the polishing run, while at the high-frequency domain (L << DH), the physics of the polishing process determine surface quality (micro-roughness). In the case of mid-spatial frequencies (L~DH) surface quality depends on the geometry and overlap of the polishing trajectories. We obtained PSD measurements as described in [4] and references therein. The PSD is discussed using the results obtained while polishing four 50 mm etalon plates to better than λ/100. These 50 mm diameter water-free fused silica plates are used in an NIR scanning Fabry-Perot interferometer. These surfaces were polished using the HyDRa tool discussed in Section 4. Surface measurements were taken with a phase-shifting (PS) Fizeau interferometer in order to quantify the figure with a 180 μm pixel size projected onto the surface. A PS Linnik interferometer with 2× and 50× objectives (equivalent pixel sizes of 7.6 and 0.16 μm) was used to determine mid- and high-spatial frequencies, respectively. In Figure 11a, the 2-dimensional power spectrum PSD2 vs. spatial frequency is plotted. It can be noted that three overlapping regimes exist that correspond to the series of instruments that were used to evaluate the surface quality. The integrated RMS values for each regime are 3.8, 1.5 and 2.9 nm for low-, mid- and high-spatial frequencies, respectively. The overall slope is approximately described as
(a) 2-D power spectral density PSD2 as a function of linear spatial frequency obtained with HyDRa on an etalon plate. (b) One of four etalon plates prior and after HyDRa polishing. The wrapped phase of the surface before (left) and after (right) polishing is shown in the upper images. Unwrapped phases of the original and polished surface, respectively (lower images). The inscribed green circle delimits the plate’s usable region. Of the 50 mm plate diameter, only the central 40 mm were polished, since the exterior ring is used for mounting the plates. This area is indicated within the light-colored circle.
Interferograms [10] showed initial figure errors that ranged between 27 and 83 nm. Using these measurements, we calculated the error maps to compute a dwell-time based raster pattern trajectory for the CNC polishing machine. An acceleration pressure of 40 PSI was chosen and tool height was controlled to achieve zero-force on the workpiece. RMS surface qualities between 3.6 and 6.8 nm were obtained after two 15 min polishing runs. The low frequency interval of the PSD shows an overall RMS fit to the desired figure of 3.8 nm, which is in accordance with the results presented in the previous section: a final surface figure quality of >>λ/100 for visible wavelengths. Sub-aperture polishing can introduce unwanted patterns associated with the polishing trajectories [14] which can occur in HyDRa polishing with a 7 mm footprint on a 40 mm sample. To minimize these mid-spatial frequencies, the tool was raster-scanned with 0.25 mm steps, which corresponds to 1/20th of the tool’s footprint size. Traces of this raster pattern should have been visible in the 0.14–4 mm−1 frequency range on the PSD. However, no evident peaks that could have been related to grooves left by a raster pattern could be observed, and only a tiny peak corresponding to about 100 μm could be noted. This peak adds a very small fraction of a nm to the total 1.5 nm RMS of the mid-frequency band, demonstrating that “over-rastering” can be effective in minimizing mid-spatial frequencies. This represents an alternative to the approach proposed by Fähnle [15], using only one tool. A comparatively large footprint has the extra advantage of making our process insensitive to CNC positioning errors, which are two orders of magnitude smaller than the footprint size. Finally, the high-spatial frequency domain shows an RMS of close to 3 nm. Although the PSD decreases in this region, it apparently stabilizes at frequencies >103 mm−1 or sizes smaller than 1 μm. This number is related to the grit size that was used in this test. To sustain a decreasing PSD tendency and thus, smaller values of the high-frequency RMS, the use of smaller grit sizes is suggested. Integrating the high-frequency domain of the PSD to obtain the RMS is also equivalent to calculating the RMS directly from a 50 × 50 μm area of the micro-interferogram, according to a standard definition of micro-roughness [16]. This micro-roughness obtained with the HyDRa process (3 nm) is comparable to the roughness reported in current FJP literature [5].
This work was funded by Universidad Nacional Autónoma de México DGAPA-PAPIIT grants IN112505, IN115509, IT100216 and IT100118, as well as by Instituto de Astronomía, UNAM.
Energy demands in the transportation sector are increasing due to a growing population and simultaneously economic policies are aiming to improve efficiency and reduce hazardous pollutant emissions including nitrogen oxides (NOx), unburned hydrocarbons (UHC) and particulate matter (PM). This has led to a great deal of interest in vehicle electrification as well as cleaner and more efficient engines. While vehicle electrification and hybridization has been growing, the cost and energy density limitations of batteries still pose challenges. As such, it is predicted that internal combustion engines will still power 60% of light-duty vehicles in 2050 [1] and the heavy-duty market will likely be mainly powered by engines for the foreseeable future.
\nIn order to abide by the stringent emissions regulations and deliver power efficiently, there is a need for clean, high efficiency engines. A variety of strategies have been investigated in order to improve the efficiency of today’s engines. These include technologies such as variable valve timing that aim to reduce pumping losses associated with the gas exchanges process and variable geometry turbochargers that seek to harness exhaust energy to improve the power density of engines. In addition, more advanced fuel injection systems have also been implemented in order to inject fuel at higher pressures and thereby promote fuel and air mixing. Improved mixing will increase the combustion efficiency and also reduce emissions of particulate matter. More complex fuel injection systems can also be used in order to develop dual-fuel combustion strategies.
\nDual-fuel combustion strategies have been demonstrated to be advantageous on both spark-ignited (SI) and compression-ignited (CI) engines. On SI engines, dual-fuel technologies can be leveraged to combat knock. Knock typically occurs in high temperature and high pressure in-cylinder conditions at which the fuel-air mixture will auto-ignite creating pressure shock waves in the cylinder. Knock can significantly damage the engine and is most prevalent at high loads where the efficiency reaches its peak. As such, high efficiency engine performance with gasoline fuel is often limited by knock. In high load conditions, the engine combustion phasing is often delayed to a suboptimal timing in order to avoid knock. While this allows harmful premature combustion to be avoided, it also leads to reductions in efficiency.
\nAlternatively, knock can also be prevented by using a fuel with a higher octane number (typically described by the research octane number (RON), motor octane number (MON) or anti-knock index (AKI)). Fuels with a high octane rating will be able to operate at the optimal combustion phasing even at high loads, but are more expensive. If high octane fuels are used in dual-fuel engines, they can enable a technique known as “octane-on-demand”. Octane-on-demand strategies are often implemented on engines with dual-fuel capabilities by using both a low RON fuel and a high RON fuel simultaneously [2, 3, 4, 5]. With dual-fuel capabilities, the fuel mixture’s knock resistance can be changed in real time to avoid knock while maintaining optimal combustion phasing. Such methods also allow fuel cost to be minimized since a less expensive, low RON fuel can be used in the lower operating conditions and the high RON fuel can be used only in knock-prone conditions.
\nOn CI engines, dual-fuel injection methods have historically been used for retrofitting old diesel engines with a cheaper fuel. In addition to the utilization of an alternative power source, the implementation also enabled reductions in PM emissions. More recently, dual-fuel injection methods have been used to promote the utilization of less reactive fuels and facilitate more advanced combustion strategies. Some dual-fuel combustion modes have shown significant promise and operate with high efficiency and low pollutant output. This is often achieved over a wide operating range by simultaneously utilizing two fuels with differing reactivities to promote premixing of the fuel or create stratification of the reactivity of the in-cylinder mixture [6, 7].
\nWhile these dual-fuel combustion modes show promise, they are not currently utilized in many production vehicles, due to a variety of challenges including difficulties with controlling combustion phasing and combustion stability with the more complex combustion strategy as well as consumer acceptance and infrastructure limitations. Currently, most of these dual-fuel combustion strategies are studied in closely monitored laboratory environments on single cylinder engines. Once removed from the laboratory and implemented on multi-cylinder engines, combustion variations and phasing challenges begin to dominate [8, 9, 10]. One such challenge is the occurrence of more significant cylinder-to-cylinder variations that can lead to inconsistent power production and potentially damaging engine conditions. In addition, on CI engines, many dual-fuel combustion strategies leverage a more premixed combustion and as such, the timing of the combustion event is controlled by the chemical kinetics. This makes it more challenging to properly time the combustion event. More advanced control methodologies are required to reduce these combustion variations and ensure an optimal combustion phasing.
\nDual-fuel engines have the potential to be highly efficient and clean, but their usage may also be limited by consumer acceptance and infrastructure challenges. Users will have to fill two fuel tanks and will need access to the needed fuels in a broad enough region. This chapter will discuss the technological developments that led to today’s dual-fuel engines, and the advancements that have been made on dual-fuel CI and SI engines.
\nThe concept of the dual-fuel engine has been around almost as long as the Gasoline (Otto) and Diesel engine. Following the development of Nikolaus Otto’s spark-ignited engine, the desire to improve the thermal efficiency by increasing the engine compression ratio led to the development of Rudolf Diesel’s compression-ignited engine. Subsequently, interest in better controlling the ignition and regulating the combustion led Rudolf Diesel, himself, to propose a dual-fuel combustion strategy and patent his invention in 1901 [11]. Today, the idea has been leveraged to promote the use of gaseous fuels such as natural gas in diesel engines and for the development of advanced combustion strategies that take advantage of the ability to dynamically optimize the properties of the fuel mixture (by controlling the proportion between the injected fuels) based on the operating conditions. Such implementations of the dual-fuel combustion strategy promise significant gains in fuel efficiency as well as reductions in toxic emissions. Nevertheless, most of the benefits associated with dual-fuel combustion have been primarily explored in academic and research institutions under strictly regulated conditions; the technology currently still faces significant challenges and limited acceptance, which restricts its market penetration.
\nThis section aims to provide an overview of the development of the dual-fuel engines by specifically reviewing the history behind the technology and discussing examples of current and past dual-fuel engines in production. The subsequent sections will discuss ongoing research on dual-fuel engines and its expected role in the near and far future.
\nIn a patent application filed on April 6, 1898, Rudolf Diesel proposes that “if a given mixture is compressed to a degree below its igniting-point, but higher than the igniting-point of a second or auxiliary combustible, then injecting this latter into the first compressed mixture will induce immediate ignition of the secondary fuel and gradual combustion of the first mixture, the combustion after ignition depending on the injection of the igniting or secondary combustible” [11]. This patent entitled Method of Igniting and Regulating Combustion for Internal Combustion Engines was accepted in 1901 and marks one of the initial efforts to introduce and successfully ignite a less reactive gaseous fuel in a 4-stroke internal combustion engine using a second fuel. Similarly, today, the ability to ignite a premixed charge (ex: air and a low reactivity fuel such as natural gas) with a secondary high reactivity fuel (such as Diesel) or interchangeably solely operating on the high reactivity fuel is one of the important characterization of a dual-fuel combustion strategy.
\nFor several years, the dual-fuel engine was not used commercially due to its mechanical complexity and rough running caused by auto-ignition and knocking. The first commercial dual-fuel engine was only produced in 1939 by the National Gas and Oil Engine Co. in Great Britain. The engine, fueled by town gas or other types of gaseous fuels, was relatively simple to operate and was mainly employed in some areas where cheap stationary power production was required [12]. During the Second World War, the shortage of liquid fuels attracted further interest in dual-fuel engines from scientists in Great Britain, Germany and Italy. Some diesel engine vehicles were successfully converted to dual-fuel and the possible application of dual-fuel engines in civil and military areas were also explored. Different kinds of gaseous fuels, such as coal gas, sewage gas or methane, were employed in conventional diesel engines during this time [13]. After the Second World War, due to economic and environmental reasons, dual-fuel engines have been further developed and employed in a very wide range of applications from stationary power production to road and marine transport, including long and short haul trucks and busses [12].
\nIn 1949, Crooks, an Engineer at The Cooper-Bessemer Corporation—one of the main engine manufacturers during World War II, presented experimental work with a dual-fuel engine that claimed to have led to the most efficient engine known with a thermal efficiency of 40% at full load. He further highlights that the dual-fuel engine has led to “an extremely economical source of power having an extremely low maintenance cost” [14]. The potential of utilizing relatively cheap gaseous fuel resources and simultaneously benefitting from high thermal efficiencies have promoted the conversion of a conventional compression ignition engine to dual fuel operation. Nevertheless, important limitations still persist: (1) at high loads, the power output and efficiency was limited by the onset of autoignition and knock with most common gaseous fuels, (2) the combustion process in dual-fuel engine is highly sensitive to the type, composition, and concentration of the gaseous fuel being used, and (3) at light load operation, the dual-fuel engine exhibits a greater degree of cyclic variations in performance parameters such as peak cylinder pressure, torque, and ignition delay [13].
\nA great deal of research is still being undertaken to understand and overcome the challenges associated with the operation of dual-fuel engines. A promising endeavor consists of successfully harnessing the benefits of the dual-fuel engine in the automotive industry.
\nIn a book chapter entitled ‘The Dual-fuel Engine’ published in 1987, Ghazi A. Karim who had previously conducted several studies [15, 16, 17, 18, 19, 20] on the topic of dual-fuel engines suggests that although dual-fuel engine has been employed in a wide range of stationary applications for power production, co-generation, compression of gases and pumping duties; the implementation in mobile applications “remain a field of urgent long term research that can have the potential for opening widely the market for the dual-fuel engine and the increased exploitation of gaseous fuel resources, particularly in the transport sector” [21].
\nIndeed, the implementation of dual-fuel technology has been more favorable in stationary and heavy-duty applications as opposed to mobile and light-duty applications. Yet, the opportune long-term research proposed by Karim for the transportation sector is still on-going. More recently, efforts to diversify the energy resources of the transportation industry have motivated researchers and engine manufacturers alike to investigate opportunities to leverage the dual-fuel combustion strategy. Furthermore, government imposed regulations on engine-out emissions and fuel efficiency targets have propelled the search for innovative engine technologies including novel implementations of the dual-fuel concept.
\nIn more recent years, a research group at the University of Wisconsin-Madison proposed the implementation of a dual-fuel combustion strategy to reduce Nitrogen Oxide (NOx) and Particulate Matter (PM) emissions [6, 7, 10, 22]. The combustion strategy called Reactivity Controlled Compression Ignition (RCCI) promises significant pollutant reductions as well as impressive fuel efficiency gains. RCCI uses in-cylinder fuel blending with at least two fuels of different reactivity and multiple injections to control in-cylinder fuel reactivity to optimize combustion phasing, duration and magnitude. The process involves introduction of a low reactivity fuel into the cylinder to create a well-mixed charge of low reactivity fuel, air and recirculated exhaust gases. The high reactivity fuel is injected before ignition of the premixed fuel occurs using single or multiple injections directly into the combustion chamber [22].
\nKokjohn et al. [6] compared the performance of a conventional diesel combustion and a dual-fuel RCCI combustion. Their study showed the implementation of a dual fuel combustion strategy yielded a reduction in NOx by three orders of magnitude, a reduction in soot by a factor of six, and an increase in gross indicated efficiency of 16.4%. Splitter et al. [7] demonstrated on a dual-fuel RCCI engine that optimizing in-cylinder fuel stratification with two fuels of large reactivity differences achieved gross indicated thermal efficiencies near 60%. Furthermore, they showed through simulations studies that a dual-fuel combustion strategy rejected less heat, and that ~94% of the maximum cycle efficiency could be achieved while simultaneously obtaining ultra-low NOx and PM emissions.
\nSimilar motivations to boost the thermal efficiency of engines have led to the implementation of a dual-fuel strategy in light duty-spark ignited engines as well. Initially proposed as an engine concept in 2005 by Cohn et al. [3], the dual-fuel spark-ignited engine featured two fuel injections systems—one for conventional gasoline and another for ethanol. The engine would promote the utilization of alternative fuels such as ethanol reducing the dependence on fossil fuels, and it was an alternative knock suppressing strategy which allows for higher load and higher efficiency operations. A high octane rating fuel such as ethanol is used in conjunction with the conventional fuel, gasoline, to dynamically adjust the fuel mixture’s resistance to auto-ignition based on the operating conditions.
\nThe studies by Cohn et al. [3] suggested dual-fuel combustion could potentially increase an SI engine’s drive cycle efficiency by approximately 30%. Similar studies by Daniel et al. [4] demonstrated that dual-injection showed benefits to the indicated efficiency and emissions at almost all loads compared to a single fuel gasoline direct injection (GDI) strategy. Furthermore, Chang et al. [77] showed a maximum 30% Well-to-Wheels (W-t-W) CO2 equivalent reduction can be achieved by utilizing a dual-fuel injection system. Numerous studies, such as [23, 24, 25], continue to explore the benefits that can be achieved through the introduction of dual-fuel combustion in the modern automotive engines.
\nIn the next sections, the application and benefits of dual-fuel combustion are separately discussed for compression-ignition and spark-ignited engines followed by concluding remarks.
\nDiesel or compression-ignition engines dominate the medium and heavy-duty markets due to their higher efficiency and high torque production capabilities. Such engines require a more reactive fuel that will auto-ignite at high pressures and temperatures. This limits the fuels that can be leveraged on CI engines. Dual-fuel engines provide a way to use less reactive fuels since they can leverage a second more reactive fuel to produce ignition. In addition, dual-fuel concepts have also been investigated as a way to reduce engine emissions. Conventional diesel combustion is diffusion controlled and is typically accompanied by high nitrogen oxide (NOx) and particulate matter (PM) emissions [26]. Nitrogen oxide emissions result from high in-cylinder temperature conditions which promote the combination of nitrogen (carried in with the fresh air) with excess oxygen [27]. Meanwhile, particulate matter or soot is produced in fuel rich regions when hydrocarbon species agglomerate [27, 28]. As such, high local equivalence ratios can lead to soot formation and high local temperatures can lead to NOx formation as shown in Figure 1. In order to avoid these problematic regions, many dual-fuel, heavy-duty CI engines attempt to operate in conditions which promote premixing of the fuel and air and/or achieve in-cylinder stratification in order to reach high efficiencies and low emissions. By enabling a more premixed combustion, rich regions where PM would be produced can be nearly eliminated and shorter combustion durations are achieved which reduces local temperatures and thereby, NOx emissions [6, 7, 29, 30, 31, 32, 33].
\nEmissions with respect to local temperature versus local equivalence ratio.
As such, dual-fuel engines have been pursued in the heavy-duty market for two main reasons:
As a way to leverage more readily available but less reactive fuels as the primary power source and use a high reactivity fuel to initiate combustion.
As a way to introduce fuels of varying reactivities and create a more complex combustion mode that can be more efficient and produce less NOx and PM.
As the world seeks to become less reliant on conventional diesel and gasoline, there has been increasing interest in using fuels such as natural gas in engines. Some of these fuels are less reactive than conventional diesel fuel and therefore, are more challenging to use on compression-ignition engines where auto-ignition of the fuel is needed. Dual-fuel systems are one way to leverage less reactive fuels on heavy-duty engines [34, 35, 36, 37, 38]. One such fuel is natural gas and it will be focused on here as an example of the benefits and challenges of this type of engine operation.
\nNatural gas is more difficult to ignite than conventional engine fuels so it is more easily integrated into spark-ignited engines. On heavy-duty engines, natural gas needs an ignition source so it is typically port-injected and diesel is direct injected and serves as a pilot. Fuels that are port-injected become premixed with the air and typically exhibit a rapid combustion event that is dominated by the chemical kinetics of the combustion reaction, but fuels that are direct injected and have to mix with the air tend to have a longer combustion event that is dominated by the time taken for the air and fuel to mix adequately. Since dual-fuel engines have a fuel that is port-injected and one that is direct-injected, they often exhibit a two-stage combustion process. The portion of combustion that occurs in a premixed vs. a diffusion mode will be strongly dependent on the amount of each fuel that is used [39]. While this makes the combustion process more complicated, dual-fuel injection can provide stable combustion of a less reactive fuel like natural gas in CI engines. However, fuel economy reductions around 10% have been observed when operating in this type of mode [34].
\nNot only is fuel economy or efficiency impacted, but emissions are also altered with dual-fuel combustion. In natural gas-diesel dual-fuel engines, up to 60% reductions in NOx and PM have been observed [34]. However, these emissions are dependent on the fuels used as well as the amounts of each fuels used. For example, particulate matter emissions and the particle size distribution of the particulates have been shown to strongly depend on the properties of the direct-injected fuel and level of natural gas substitution. Direct injected fuels with lower densities and viscosity and higher volatility produce lower amounts of particulates [40]. However, higher natural gas substitution rates can increase soot levels since they decrease the local oxygen availability [41].
\nAs with many natural gas engines, higher CO and UHC emissions are typically encountered. Various natural gas substitution rates have been explored in [42] and showed that only lower amounts of natural gas could be used at low load conditions due to emissions constraints, but higher fractions of natural gas could be used at high loads. Direct injection of both fuels [43], higher fuel injection pressures, and adapted engine control units [44, 45] have been implemented to avoid these emissions constraints. After treatment systems including diesel oxidation catalysts [35] as well as diesel particulate filters and urea-selective catalytic reduction systems [46] have also been introduced on dual-fuel engines to reduce emissions. However, to enable efficient use of high amounts of natural gas, more advanced combustion methods and optimization methods are likely needed [47, 48].
\nA majority of conventional dual-fuel engine studies have focused on natural gas, but this approach of using diesel as a pilot fuel can also be leveraged with a variety of fuels that are not reactive enough to be used as the sole fuel on a compression-ignition engine. Dual-fuel concepts have also been explored with fuel combinations including on methanol and diesel [49], biogas and biodiesel and biogas and diesel [50].
\nIn order to push engines to higher efficiencies, there has been a great deal of exploration into more complex combustion modes. Many of these advanced combustion strategies attempted to premix the fuel and air in order to achieve a more efficient and clean combustion, but were only able to be leveraged in lower torque ranges [51, 52]. One strategy for expanding the operating region of these more advanced techniques is to simultaneously utilize two fuels with differing reactivities in order to further increase the combustion delay period and promote premixing in higher operating regions [53]. This strategy is known as reactivity-controlled compression ignition (RCCI). In RCCI, a fuel with low reactivity such as gasoline is injected separately from a high reactivity diesel-type fuel. The quantities of each respective fuel can be modified so that the combustion event can be delayed to provide adequate mixing time and the desired shape of the combustion event can be achieved. Recent work in RCCI has shown that fuel properties that differ from those of conventional fuels can be leveraged to shape the combustion process and increase engine efficiency from 45% to near 60% [6, 7] in this mode. While the efficiency benefits can be significant, high CO and UHC emissions as well as high pressure rise rates can still limit the use of RCCI.
\nRCCI-type combustion was originally studied at the University of Wisconsin-Madison using gasoline as the port-injected low reactivity fuel and diesel as the direct-injected, high reactivity fuel [7]. By leveraging two fuels with varying properties stratification of the in-cylinder mixture reactivity could be achieved leading to longer ignition delays and increased time for premixing. Diesel fuels with lower reactivities were shown to be advantageous in these operating conditions as they increase the local reactivity gradient [54, 55]. In such modes, the more reactive fuel components are consumed at a faster rate and the slower burning competent make up a larger portion of the UHC emissions [56].
\nThe use of alternative fuels such as ethanol and natural gas in such RCCI-type operation conditions has also shown promising results and appears to better take advantage of these alternatives. Research by Navistar, Argonne National Laboratory and Wisconsin Engine Research Consultants found that using E85 as the low reactivity fuel could allow higher loads and efficiencies to be achieved with RCCI. While more traditional gasoline and diesel dual-fuel operation achieved a BMEP of 11.6 bar and brake thermal efficiency (BTE) of 43.6%, using E85 with diesel allowed operation to be extended to 19 bar BMEP with a BTE of 45.1% [57]. Later studies led by RWTH Aachen University and FEV GmbH showed that when using diesel and ethanol, higher ethanol quantities could be leveraged in lower load conditions and would provide a more stable combustion and lower UHC emissions. However, as the load was increased higher amounts of diesel were required in order to keep the cylinder pressure rise rate to an acceptable level [58].
\nSome of these detrimental impacts on CO and UHC emissions are able to be counteracted by more complex fuel injection strategies [59]. For example, more recent work has explored the use of ethanol port-injection with a multi-pulse direct-injection of diesel. A double pilot injection was able to reduce the UHC and CO emissions in RCCI-type conditions [60]. Other methods such as leveraging higher injection pressures have also been shown to be able to increase efficiency and provide further decreases in emissions of NOx, CO, UHC, and PM [61].
\nWhile RCCI methods are promising, these modes suffer from several technical challenges. First, cycle-to-cycle and cylinder-to-cylinder variations can be more dramatic than in conventional diesel combustion [62, 63]. Since fuel and air mixing are critical and high amounts of recirculated exhaust gas are typically leveraged in these modes, small variations in the in-cylinder fuel quantities and gas mixture can lead to significant variations in the combustion process. Combating such variations is likely to require more complex control strategies and additional engine sensors [62].
\nSecond, control of the combustion phasing of these modes is challenging since the combustion process is controlled by chemical kinetics and not directly triggered by an injection event. Control techniques must try to maintain an optimal combustion phasing while ensuring that pressure rise rate and combustion variations do not exceed acceptable limits by monitoring the fuel blend ratio [64] and direct-injection timing [65]. Successful use of RCCI may also require switching between traditional diesel combustion at lower load and dual-fuel operation at higher load [65]. Intermediate modes such as “premixed dual-fuel combustion” and “partially premixed compression ignition” may provide clean and efficient intermediate combustion strategies that can be used on their own or in transitions from conventional diesel combustion to RCCI [66].
\nAs discussed previously, UHC and CO emissions are often higher in RCCI modes. This is believed to be because local equivalence ratios can drop below the flammability limit of natural gas and lead to unburned hydrocarbon emissions [67], but may necessitate the development of new after treatment systems for these engines. Consumer acceptance is also a concern with dual-fuel engines. Since users may not want to fill two fuel tanks, Splitter et al. explored a method of enabling RCCI by using gasoline and the cetane improver di-tert-dutyl peroide (DTBP) [22]. This study leveraged port injected gasoline as the low reactivity fuel, but used gasoline mixed with varying amounts of DTBP as the high reactivity fuel. A peak gross ITE of 57% was achieved and emissions were similar to standard dual-fuel RCCI levels.
\nThe implementation of dual-fuel combustion strategies on medium and heavy-duty engines have primarily been discussed in the preceding section. These utilizations of dual-fuel combustion strategies are generally restricted to compression-ignition engines where Diesel is the conventional fuel. Nevertheless, in recent years, light-duty spark ignited engines have also been configured to feature a dual-fuel combustion system. This section will discuss the implementation and utilization of dual-fuel combustion on light-duty SI engines.
\nIn SI engines, a dual-fuel combustion strategy is also leveraged to promote the utilization of alternative fuels such as ethanol and methanol. There are increasingly numerous government imposed legislations promoting the use of biofuels in transportation [68]. Current legislation requires EU member states to conform to a 10% minimum target on the use of alternative fuels (biofuels or other renewable fuels) in transportation by 2020 [69] . In the US, tax incentives have been used to promote the use of ethanol in gasoline [70], in order to replicate the success seen in Brazil [71]. The quest to benefit from incentives or conform to legislations has led engine manufacturers to explore the implementation of dual-fuel combustion on spark-ignited engines with the utilization of biofuels and other renewable fuels.
\nAdditionally, one of the primary motivations for the implementation of dual-fuel combustion in SI engines has also been for the development of better engine knock control techniques. Engine knock, the inadvertent auto-ignition of the fuel in localized high pressure and temperature regions inside the cylinder [72, 73], can result in significant engine damage and marks one of the main obstacles in SI engines. The conventional approach to avoiding knock in spark-ignited (SI) engines consists of delaying the combustion phasing by retarding the spark timing [74]. A combustion event occurring later in the combustion stroke (further away from top dead center) has a lower tendency to knock since the combustion pressure and temperature are lower. However, delaying the combustion phasing also reduces the fuel efficiency since less work can be extracted by the late combustion [75].
\nA dual-fuel combustion strategy provides SI engines with an alternative way to avoid knock without sacrificing fuel efficiency. The tendency of the fuel to auto-ignite is not only dependent on the in-cylinder conditions, but also on the knock resistance (octane rating) of the fuel. As such, increasing the fuel’s octane rating helps avoid knock without compromising fuel efficiency. Engines with dual-fuel capabilities can use a low RON fuel and a high RON fuel simultaneously to optimize the fuel mixture’s knock resistance by controlling the proportion of each injected fuel. Many studies have explored the implementation of a dual-fuel strategy to suppress knock [2, 3, 4, 5, 72, 73, 74, 75].
\nThe studies by Cohn et al. [3] and Bromberg et al. [76] at Massachusetts Institute of Technology proposed an ethanol boosted engine concept, which provides suppression of engine knock at high pressure through the use of direct ethanol injection. Their studies conclude that the implementation of the secondary fuel injection system could allow engine operation at much higher levels of turbocharging and could potentially increase the drive cycle efficiency by approximately 30%. Daniel et al. [4] implemented a dual-fuel strategy for knock mitigation on a single cylinder SI research engine. The study shows that dual-injection strategy (using either ethanol or methanol as the high RON fuels) showed benefits to the indicated efficiency and emissions (HC, CO, CO2) at almost all loads compared to a single fuel gasoline direct injection (GDI) strategy. Furthermore, Chang et al. [77] conducted a Well-to-Wheels (W-t-W) greenhouse gas emissions assessment to estimate the overall emissions benefits of a knock mitigating dual-fuel system. Their study showed a maximum 30% W-t-W CO2 equivalent reduction can be achieved by utilizing a dual-fuel injection system.
\nA dual-fuel SI configuration provides three main benefits: (1) engine can be further downsized and operated in high pressure conditions (2) the fuel knock resistance can be adjusted based on operating point while maintaining an optimal combustion phasing (maximizing engine efficiency), and (3) operating points with low knock propensity can be operated with a low octane fuel, eliminating the waste of RON, which generally translates into cheaper fuel cost and lower CO2 emissions [5, 77]. A team at Saudi Aramco, in collaboration with IFP Energies nouvelles, demonstrated these benefits on a production passenger vehicle [23, 78, 79]. The dual-fuel technology is identified as “an opportunity to improve fuel efficiency by using the octane only when you need it.” The researchers outline that the technology will improve fuel efficiency while reducing overall energy requirements to manufacture gasoline fuels in the future [79]. The researchers at Aramco Fuel Research Center (AFRC) identify the development of the production car as only the start, and outline the near-term objective of going from a vehicle with two tanks with two different fuels to one that only uses only one fuel and process it with an on-board fuel upgrading system.
\nSimilar to the efforts of Saudi Aramco, there is currently a growing interest to harness the benefits of a dual-fuel combustion strategy on conventional SI engines. A study at Massachusetts Institute of Technology by Jo et al. [24] investigates the use of dual-fuel for a passenger vehicle and a medium-duty truck. Their simulation studies, coupled with experimental testing, conclude that significant gains in engine brake efficiencies can be achieved: 30% for the Urban Dynamometer Driving Schedule (UDDS) cycle and 15% for the US06 cycle. Studies by Marchitto et al. at Istituto Motori, CNR [25] demonstrate the port injection of ethanol as a secondary fuel showed a significant increase in thermal efficiency (~10%) and significant reduction in particle number emissions as well as particulate mass (60–80%).
\nDual-fuel combustion provides a promising path to boost the thermal efficiency of spark-ignited engines. The opportunity to leverage alternative fuels such as ethanol and methanol, as well as the ability to suppress knock without compromising thermal efficiency, has garnered interest in the technology. Nevertheless, the challenges associated with multiple tanks and multiple fuels require researchers to seek paths that will make the technology more accessible to everyday consumers.
\nSince the inception of the dual-fuel combustion strategy as a tool to better control combustion, its application has been most vital in stationary and heavy-duty applications. The integration of the dual-fuel combustion in the transportation industry promises great benefits both in terms of fuel efficiency improvement as well as toxic emissions reduction. Significant efforts are undertaken to implement this technology in the automotive industry for heavy-duty, as well as medium and light-duty engines. The on-going research on dual-fuel combustion promises encouraging paths that will allow the utilization of more readily available gaseous fuels and renewable fuels. The observed benefits on both compression-ignition and spark-ignited engines warrants further investments, research, and efforts to better exploit these gains on a larger scale.
\nIn compression ignition (CI) engines, dual-fuel combustion presents an effective approach to control combustion timing and extend engine load limitations. This is achieved by injecting both a high reactivity and low reactivity fuel, adjusting the concentration of one relative to the other, and thereby optimizing the fuel mixture’s reactivity (on a cycle-by-cycle basis) for different operating conditions. In spark ignited (SI) engines, the optimization of the fuel’s octane rating presents an alternative approach to avoid abnormal combustion (engine knock due to auto-ignition). The conventional SI engine relies on delayed spark timings to avoid knock, which results in degraded efficiency and higher emissions. With dual-fuel combustion techniques, two fuels with high and low octane ratings can be used to adjust the fuel mixture’s octane as needed to avoid knock without sacrificing the engine’s efficiency.
\nWhile dual-fuel operation has many potential benefits, the implementation of these strategies on CI and SI engines also involve significant challenges. These challenges are exacerbated when dual-fuel combustion is implemented in conjunction with other advanced combustion strategies including varying valve timing, and high EGR circulation. The underlying challenges include increased combustion variations and difficulties in properly adjusting fuel mixture for effective control of combustion timing (in CI engines) and effective knock control (in SI engines). While the technology has successfully been used in stationary applications, the implementation of dual-fuel strategy on mobile applications, specifically in the transportation sector, still faces limiting challenges. In addition to the technical challenges associated with the dual-fuel engine, a primary concern for its integration in the automotive industry consists of the social resistance to the requirement of having and filling two fuel tanks. In order for the technology to successfully penetrate the automotive market, the benefit in the terms of fuel efficiency improvement and toxic emissions reduction need to clearly outweigh the technical and social challenges.
\nYuri Gagarin was the first human to travel into space in 1961, and subsequently, more than 500 people have followed. We have learned so much since then, however, there are many questions that remain unanswered. The Humans in Space program seeks to facilitate the creation of new knowledge that will help find solutions to these questions through the development of a first-of-its-kind and more complete library of open access space life sciences and related areas knowledge.
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\n\nThe mission of the Humans in Space Program is to create a unique, open-access reservoir of knowledge in the area of space life science, based on an interdisciplinary and comprehensive approach that will promote interaction between world-leading experts, academics and researchers. With education and the global dissemination of research at its core, the program will be an ambitious collection of experts, high-quality literature and lectures.
\n\nWithin the first 5 years, the program will:
\n\nDrawing from the extensive network of space scientists, academics, researchers, collaborators and mentors associated with its partners, the program will lay the foundations for a web of knowledge that will grow and nurture future generations of scientists, engineers, thought leaders and explorers.
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