cutting tool edge engineering

Wayan Darmawan

CUTTING TOOL EDGE ENGINEERING FOR ECHO-MACHINING OF WOOD

CUTTING TOOL EDGEENGINEERING FOR ECHO-MACHINING OF WOOD

PT Penerbit IPB Press

IPB Science Techno Park Jl. Taman Kencana No. 3, Bogor 16128 Telp. 0251 - 8355 158 E-mail: [email protected] Penerbit IPB Press

@IPBpress


xxxxxxxxx ISBN : 978-602-440-xxx-x

Wayan Darmawan



Wayan Darmawan

Penerbit IPB Press IPB Science Techno Park, Kota Bogor - Indonesia

C.01/06.2017

Judul Buku: Cutting Tool Edge Engineering for Echo-Machining of Wood Penulis: Wayan Darmawan Editor: Helda Astika Siregar Desain Sampul & Penata Isi: Ahmad Syahrul Fakhri Korektor: Nopionna Dwi Andari Jumlah Halaman: 122 + 10 halaman romawi Edisi/Cetakan: Cetakan 1, Juni 2017 Sumber Illustrasi Sampul: http://www.emuge.com/news/end-mils PT Penerbit IPB Press Anggota IKAPI IPB Science Techno Park Jl. Taman Kencana No. 3, Bogor 16128 Telp. 0251 - 8355 158 E-mail: [email protected] ISBN: 978-602-440-109-2 Dicetak oleh Percetakan IPB, Bogor - Indonesia Isi di Luar Tanggung Jawab Percetakan © 2017, HAK CIPTA DILINDUNGI OLEH UNDANG-UNDANG Dilarang mengutip atau memperbanyak sebagian atau seluruh isi buku tanpa izin tertulis dari penerbit

Prologue

Research topics on wood machining are very broad. Separate reports, which cover a special subject of the field, exist. For example, there are reports dealing with wood cutting mechanisms, monitoring and controlling cutting processes, cutting forces, cutting noise, and tool wear characteristics in cutting wood, etc. The contents described in this book deal with the wear, force, noise, and surface characteristics in machining of wood and wood-based materials using engineered cutting tools. The contents are presented in eight chapters. Introduction which consists of the importance of surface engineering in broad sectors, and the development of cutting tools for wood and wood-based machining are cited in chapter I, and the lists of references are also included. Part II is a discussion of experimental results dealing with the importance of extractives and abrasives in wood and wood-based material on the chemical and mechanical wearing of high speed steel and tungsten carbide cutting tools. Surface engineering on the cutting tools edge for environmentally friendly machining of wood and wood-based materials is discussed in chapter III to VIII. Chapter III discusses the quality (the wear, force and noise characteristics) of the monolayer coated cutting tools in machining high-density hardboard. Chapter IV deals with the performance (wear, forces, and noise characteristics) of the monolayer coated cutting tools for high speed machining of wood-chip cement board. Chapter V discusses on wear and delamination resistance, and wear mechanisms of multilayer coated cutting tools in machining wood and wood-based materials. Chapter VI deals with improvement of cutting performance of high speed steel cutting tool by surface modification of laser melting. Chapter VII discusses the cutting performance of high speed steel cutting tool by surface modification of laser cladding. Chapter VIII discusses the development of new design of extreme helical edge milling cutter for planning wood.

Cutting Tool Edge Engineering for Echo-Machining of Wood

Literature references, which are cited in the texts, are listed at the end of each chapter, allowing the reader to seek further specialized information. The total contents of this book represent the expantion of our articles published at peer reviewed journals. My sincerest thanks are due to colleagues (Dodi Nandika, Irsan Alipraja, Fauzan Fahrussiam, Kadiman, Suhada) for their kind assistances. I am particularly thankful to Prof. emeritus Chiaki TANAKA, Dr Hiroshi USUKI (Faculty of Science and Engineering, Shimane University), who encouraged me for the work on the coated cutting tool), Prof. Remy MARCHAL and Dr Jean QUESADA for their cooperation during sabbatical work at the LABOMAP ENSAM Cluny), Prof. André Wagenführ and Christian Gottlöber for their joint research work during sabbatical work at the Institute of Wood and Fibre Material Technology, Technische Universität Dresden. PT Kanefusa Japan (Dr S NISHIO and Mr Indra MALELA for the coated cutting tools and companies making wood-based materials are acknowledged. Special acknowledgment is also made to the the Directorate for Research and Community Service of the Ministry of National Education for the Republic of Indonesia for the research grant and research mobility. Acknowledgment is also made to publisher (IPB Press for permission to produce copy right material).

Wayan Darmawan

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Contents

Prologue.......................................................................................................v Contents....................................................................................................vii I

Introduction....................................................................................... 1

1.1 Surface Engineering.................................................................... 1

1.2 Wood and Wood-based Material Machining.............................. 5

1.3 Cutting Tools for Wood and Wood-Based Machining................ 6

1.4 Tool Wear in Wood and Wood-Based Machining...................... 8

1.5 Improvement in Wear Resistance of Cutting Tools................... 12

II

The Importance of Extractives and Abrasives in Wood Materials on the Wearing of Cutting Tools...................................... 19

2.1 Chemical Wear Caused by Extractive........................................ 20

2.2 Mechanical Wear of Tool Bits................................................... 23

2.3 Importance of Silica and Extractive in Wearing of Cutting Tools........................................................................... 29 III Machining Characteristics of Monolayer Coated Tools in Cutting Hardboard.......................................................................... 33

3.1 Wear and Delamination Characteristics.................................... 33

3.2 Effect of Tool Material Hardness on the Wear and Delamination............................................................ 37


3.3 Force and Noise Characteristics of the Coated Tools................. 39

3.4 Effect of Work Materials on the Wear and Delamination.......... 42

IV. Performance of Monolayer Coated Tools in High Speed Cutting of Wood-Chip Cement Board......................... 51

4.1 Effect of Work Materials on Cutting Temperature.................... 51

4.2 Clearance Wear and Cutting Tools Life ................................... 54

4.3 The Effect of Cutting Speed on Forces and Noice..................... 61

V.

Wear Characteristics of Multilayer-Coated Cutting Tools in Milling Particleboard.................................................................... 69

5.1 Delamination Wear................................................................... 69

5.2 Wear Mechanism...................................................................... 74

VI Characteristics of Laser-Melted T1 High Speed Steel and Its Wear Resistance.................................................................... 80

6.1 Effect of Laser Treatments on Depth of Melted Zone............... 81

6.2 Miscrostructures of the Melted Zone (MZ)............................... 83

6.3 The Effect of Laser Treatment on the Hardness......................... 86

6.4 Wear Resistance of the Laser Melted T1 Cutting Tool.............. 90

VII Characteristics of Laser-Cladding M2 High Speed Steel Cutting Tools................................................................................... 95

7.1 Microstructures of Laser Treated M2........................................ 98

7.2 Microhardness of the Laser Treated M2.................................... 99

7.3 Wear Resistance of the Laser Treated M2 Cutting Tool.......... 100

7.4 Behaviour of Parallel Cutting Force (Vertical force)................ 103

7.5 Roughness (Ra) Value of the Veneer Surfaces.......................... 104

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Contents

VIII New Design of Helical Edge Milling Cutters for Planing Wood..... 107

8.1 Chip Flow............................................................................... 111

8.2 Cutting Power Consumption.................................................. 113

8.3 Noise Emission....................................................................... 114

8.4 Wear Resistance...................................................................... 115

8.5 Roughness (Ra) Value of the Planed Lumber.......................... 117

About the Author.................................................................................... 121

ix

I. Introduction 1.1 Surface Engineering SURFACE ENGINEERING is a multidisciplinary activity intended to tailor the properties of the surfaces of engineering components so that their function and serviceability can be improved (Figure 1.1). Surface engineering refers to a wide range of technologies that aim to design and modify the surface properties of components. There are three categories of surface engineering methods that can be used to optimise the surface properties and the bulk materials. These are surface coatings, surface modification, and redesigning the surface shape.

Enngineering M Materials

Figure 1.1 The surface engineering road map: a multisectoral, interdisciplinary technology


1.1.1 Surface Coating Surface coating processes involve depositing a layer of molten, semi-molten or chemical material onto a substrate (Figure 1.2). One of the main functions of surface coating is to modify and reinforce the surface functions instead of reforming the composition of the bulk material. Some examples of surface coating processes include Physical Vapour Deposition (PVD), Chemical Vapour Deposition (CVD), plasma and thermal spraying, sol-gel, cladding, and electroplating [1]. Surface modification processes can be classified as hardening by flame, induction, laser or electron beam, high energy treatments, and diffusion treatments (Figure 1.2). Surface modification processes are applicable to control friction, improve surface wear and corrosion resistance, and change the physical or mechanical properties of the component. Surface modification treatments also can be combined with surface coating processes, for instance laser cladding. This combination enhances the advantages of surface coatings and surface modification, thus achieving specific requirements and fitness for purpose.

Figure 1.2 Technology classification on surface engineering The application of PVD techniques ranges over a wide variety of applications from decorative, to high temperature superconducting films. A very large number of inorganic materials-metals, alloys, compound, and mixtures as well as some organic materials, like polymers, can be deposited using PVD technologies. Nowadays, PVD is used to form multilayer coatings, gradient-

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I. Introduction

depositions, or very thick deposits [2]. PVD processes include a wide range of vaporphase technologies. In general, PVD is used to describe variety of methods to deposit thin solid films by the condensation of a vaporized form of the solid material onto various surfaces. Namely, PVD process involves physical ejection of material as atoms or molecules and condensation and nucleation of these atoms onto a substrate. The vaporphase material can consist of ions or plasma and is often chemically reacted with gases introduced into the vapor, called reactive deposition, to form new compounds. The thicknesses of deposited layers could be from few nanometers to thousands of nm [3]. There are other thin film deposition technologies called nonPVD deposition technologies that can be used for particular applications. They are: a) Chemical Vapor Deposition (CVD) and plasma enhanced CVD (PECVD); b) Electroplating, Electroless Plating, and Displacement Plating; c) Chemical Reduction. Plasma-based chemical vapor deposition involves chemical reactions on the substrate surface and most likely reactions in the gas phase as well. The treatment temperature in this process is usually higher in order to enable chemical reactions. In general, this process allows the deposition of metals, alloys, ceramic, and polymer thin films.

1.1.2 Surface Modification Modification of surface properties over multiple length scales plays an important role in optimizing a material’s performance for a given application. For instance, the cosmetic appearance of a surface and its absorption properties can be controlled by altering its texture [4,5] and presence of chemical impurities in the surface [6]. A material’s susceptibility to wear and surface damage can be reduced by altering its surface chemistry, morphology, and crystal structure [7]. Also, one can consider the frictional, adhesive, and wetting forces acting at a material interface as being strongly influenced by the size and shape of the micro and nanoscale features present [8]. As such, multiscale surface modifications are a critical factor in the development of new material structures and in engineering the detailed interactions that occur at surfaces and interfaces. In recent years, surface modification using advanced heat source like laser has been replacing the conventional methods to produce amorphous microstructure via rapid solidification. Due to the benefits of laser to enhance

3


the tribological and mechanical properties of materials’ surface, several laser surface processing were developed including laser surface modification, namely laser alloying, laser melting, and laser cladding. In high temperature applications, the laser surface modification technique is beneficial to prolong the die life cycle, and also to improve the surface roughness of thermal barrier coatings (TBC). To produce the amorphous layer at a particular depth, laser parameter such as irradiance, frequency, and exposure time are controlled. Variations of parameter may result in modified microhardness properties of heat affected zone and transition zone. Nevertheless, works on laser glazing of bearings, railroad rails and TBC had proven the surface properties were enhanced through laser glazing to cope with excessive load, wear, fatigue, bending and friction deman. Currently, high power lasers have become increasingly accepted as tools for many applications from cutting, to surface modification methods. Laser has also been proven to be capable of producing adherent, hard, wear, corrosion, fatigue and fracture resistant coatings on a diverse range of materials [9,10]. In other words, the crystal structure of metals’ surface can actually be modified into very fine non equilibrium microstructures as a result of rapid solidification (106–1012K/s) via laser surface modification [11].

1.1.3 Redesigning Surface Shape In the context of wood machining, a cutting tool or cutter is any tool that is used to remove material from the workpiece by means of shear deformation. Cutting may be accomplished by single-point or multipoint tools. Singlepoint tools are used in turning, shaping, planing and similar operations, and remove material by means of one cutting edge. Milling and drilling tools are often multipoint tools. Grinding tools are also multipoint tools. Each grain of abrasive functions as a microscopic single-point cutting edge (although of high negative rake angle), and shears a tiny chip. Cutting tools must be made of a material harder than the material which is to be cut, and the tool must be able to withstand the heat generated in the metalcutting process. Also, the tool must have a specific geometry, with clearance angles designed so that the cutting edge can contact the workpiece without the rest of the tool dragging on the workpiece surface. The angle of the cutting

4

I. Introduction

face is also important, as is the flute width, number of flutes or teeth, and margin size. In order to have a long working life, all of the above must be optimized, plus the speeds and feeds at which the tool is run. Optimisation and improvements in wood cutting operations are characterized by various approaches, including design of cutting tool, selection of the cutting tool material, and application of machining conditions. All of the approaches lead to much higher productivity, more economical cutting and reduction of the overall machining cost, in which resulted from better efficiency, stability, accuracy, and tool life during the cutting processes. For economical and high performance cutting, all parts involved in the cutting processes should be selected and optimised. Among the parts, design of cutting tool edge involved in the cutting processes would be very important.

1.2 Wood and Wood-based Material Machining Tropical wood and wood-based materials are processed in large and increasing quantities in many countries. Production of the wood-based materials was reported to be 250 and 80 million m3, respectively for Asia-Pacific region and Europe in 2015 [12]. Recently, the use of the wood and wood materials has been increasing for building constructions and decorative purposes. In the secondary wood manufacturing industry, where wood and wood-based materials such as particleboard and fiberboard are machined extensively, tool wear becomes an important economic parameter. Both woods and wood-based materials can be machined by cutting tools used in wood working industry. Consequently, no special treatment or technique is required. However, it should be considered that wood-based materials consist of wood, artificial resin and additives of different kinds. In general these materials are harder, dryer and more abrasive than solid wood. Therefore, the machining characteristics of the cutting tools in cutting wood-based materials would be significantly different from that of solid wood. Since 1930’s the use of tungsten carbide cutting tools in the wood working industry has provided significant improvements in tool life, primarily because of the superior hardness of these alloys compared to that of carbon steels, tool steels, high speed steels (HSS) and stellite (cobalt-based alloy cutting tool).

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The improvement in tool life has been particularly significant in machining abrasive materials, such as silica-containing tropical wood species and some wood-based materials. Though tungsten carbide tools have been widely used for machining solid wood, however a need for carbide cutting tools with longer life and better performance exists especially for cutting wood-based materials. An alternative, application of hard coating materials by using chemical vapor deposition (CVD) and physical vapor deposition (PVD) methods onto the surface of the tungsten carbide tools has been recently promoted for cutting wood-based materials. The use of coated carbide tools in the field of metal machining has been already well known, but has not been applied yet in the field of wood and wood-based materials machining. The scope of the present research is centered to exploring the applicability of the available coated carbide tools for wood and wood-based materials machining. Some experiments have been performed for the coated carbide tools in cutting wood and wood-based materials. The information provided for machining behaviors should be discussed for promoting their applications especially for high speed machining of difficult-to-cut wood-based materials, such as woodchip cement board and hardboard.

1.3 Cutting Tools for Wood and Wood-Based Machining During World War I high speed steel (HSS) was developed and used extensively in the metal working industry. Shortly after the war, it began to find its place in wood working industry. At present the HSS is still manufactured for cutting tools in the wood working industry. HSS is a common cutting tool material for planer blades, moulder and shaper knives, and router bits. The most important property of all HSS is its retention of hardness at elevated temperature [13]. Therefore, HSS is a good selection for many wood machining processes, as high temperature exists at the tool edge and it accelerates the tool wear [14,15]. There are many different grades of the HSS cutting tool. However, it was noted that M-2 grade of HSS cutting tool provided better performance than the other grades for machining woods due to its good toughness and excellent wear resistance [16].

6

I. Introduction

Since World War II, there was a great deal of efforts by the knife manufactures to find a higher quality wood cutting tool. Most plants have tried tungsten carbide tools, and have obtained good results. It was noted that tungsten carbide tool has lasted about 20 to 50 times longer than steel cutting tools for cutting wood [17]. Therefore, at present, tungsten carbide tools are dominating other cutting tools in the wood working industry. They are used for tipped circular saws, router bits, profile cutters, and other purposes. They have a wide range of toughness and hardness, which explain their wide use. The main constituents of cemented tungsten carbides are tungsten carbide (WC), and a metallic binder, which is mostly cobalt [18]. The manufacturing process of tungsten carbide tools allows them to be produced in several insert shapes, on which complex cutting edge profiles can be ground more easily with a diamond wheel. These inserts can either be brazed or clamped on the tool holder. Such flexibility in insert design combined with a wide range of mechanical properties allows the use of tungsten carbide tools in the economical machining of wood [19]. Another cutting tool material that had captured the imagination of the manufacturers in the metal working industry is ceramic cutting tool. The first ceramic cutting tool inserts were put on the market in 1956 [17]. Ceramic cutting tools are being used successfully in the metal machining. In some operations, ceramic tool bits can be operated at approximately twice the speed of tungsten carbide tools. However, several manufactures of the ceramic tips stated that the useful life of ceramic cutting tools for wood machining are exactly the same as that of tungsten carbide cutting tools [17]. It was also noted in another study that the wear resistance of the ceramic cutting tool was observed to be same as that of the tungsten carbide tool during machining of particleboard [20,21,22]. Furthermore, ceramic cutting tools are brittle, and its initial cost is from 40 to 200% higher than that of tungsten carbide tools [23]. Because of these drawbacks, the use of ceramic cutting tools is limited in the wood working industry. Cubic boron nitride (CBN) is a cutting tool material that is much harder than the tungsten carbide tools. Its high performance for cutting hard metals at high speed was reported to be due to the retention of strength at higher temperatures compared to tungsten carbide tools [24]. It was noted in another study [20] that the CBN cutting tool was observed to be superior in wear resistance than the tungsten carbide, cermet, and ceramic cutting tool, and 7


was found to be the same in performance as the artificial diamond cutting tool in cutting particleboard. However, the cost of the CBN cutting tool seems to remain very high for the wood machining application. Another cutting tool material that is slightly higher in hardness than the CBN cutting tool is polycrystalline diamond (PCD) cutting tool. The hardness and wear resistance of the PCD cutting tool are much better than those of tungsten carbide tools. However, its toughness is somewhat less than that of tungsten carbide tools. PCD cutting tool is commonly used as circular saw blade tips for machining wood-based materials, such as particleboard and fiberboard. The life of the PCD cutting tool was reported to be longer than that of tungsten carbide tool during cutting wood and wood-based materials [25,26]. However, it was reported that tipping of saw blades with PCD tips is not cost-effective for sawing incense cedar wood primarily due to chipping of the saw tips [27]. Because of technical problems in the making of PCD tool, such as the difficulty in creating tools of complex shapes and tool size limits, the PCD tools are available in practical use of wood-based materials machining, such as particleboard and hardboard.

1.4 Tool Wear in Wood and Wood-Based Machining The elucidation of the wear mechanisms of cutting tools is considered to be necessary in an effort to find better choices of tool materials and improve wear resistance of tool materials. During machining wood and wood-based materials, several wear mechanisms may simultaneously happen to the cutting tools. Chemical and electrochemical corrosive wear of tungsten carbide tools when machining green wood was reported to take place by preferential dissolution of cobalt binder through chemical attack by the extractives present in green wood [28,29,30,31]. It was also noted that resistance to corrosive wear is increased by adding chromium and nickel to the cobalt binder, by decreasing cobalt in content, and by decreasing grain size of the tungsten carbide [28]. In another study [32], the wear resistance of stellite cutting tool (59% Co, 29% Cr, 9% WC and 3% C), in which cobalt is alloyed with chromium, was compared to that of tungsten carbide tips for sawing unseasoned red cedar.

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I. Introduction

It was noted in the study that the stellite cutting tool was more resistant to corrosive wear than tungsten carbide tool with high percentage of cobalt in content. It was demonstrated in literatures [33,34] that electrochemical wear of tungsten carbide tool can be reduced by applying a negative electric potential between the cutting tool and the work material. By comparing edge wear for tools with and without negative electric potential, it was able to separate the mechanical and corrosive components of tool wear. It was noted in the study that corrosive wear constitutes about 40 to 65% of the total wear, and that an increase of cobalt in content and grain size leads generally proportional increase in corrosive wear. Corrosive wear characteristics of tungsten carbide tool and PCD tool were examined in cutting air-dried and wet western red cedar in another study [31]. It was found in the study that tungsten carbide and PCD tools exhibited marked corrosive wear in cutting wet western red cedar. The presence of sulfur, chlorine, and phosphorus was detected on the wear surface of tool by energy dispersive spectroscopy (EDS) and was attributed to thermal decomposition of medium density fiberboard (MDF) at temperature about 500 and 700oC [35]. Increasing temperature was reported to produce a more significant effect on the rate of oxidation. It was also noted in the work that oxidation of tungsten carbide tool starts at temperatures as low as 500oC, and temperature over 1000oC have been estimated to occur during certain MDF cutting processes [36]. Another study [37] gives an indication that strong bases (e.g. ammonium chloride and sodium chloride) were found to cause much higher normal force and tool wear than ammonium sulfate and sodium sulfate used in urea formaldehyde (UF) resins in particleboard manufacturing. Five grades of tungsten carbide were experimented in continuous and interrupted cutting of particleboard [38,39]. The grain sizes and cobalt contents were varied from 0.7 µm to 1.7 µm and 3.0% to 9.5%, respectively. It was noted in the study that the wear resistances of the tungsten carbide tool with finer grain size and lower cobalt content are better than that of coarser grain sizes and higher cobalt content. It was also noted that the wear of the carbide tools in machining the particleboard occurred primarily on the clearance face of the tool and was caused by preferential removal of the

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binder phase between the tungsten carbide grains. The binder removal caused the carbide grains to break into small fragments, which were subsequently removed. The wear resistance of the carbide tool is well correlated with the hardness of the tool material. It was also noted in the work that similarities of wear characteristics of the tungsten carbide grades tested on lathe and router suggest that wear of the tungsten carbide tools was caused by micro-abrasion and extrusion of the binder phase. Performance of different grades of tungsten carbide tools was investigated during the peripheral milling of particleboard in another literature [40]. The grain sizes of the tungsten carbide tools were varied from 0.5 to 2.0 µm, and their hardness was varied from 1750 to 2350 Hv. It was noted in the study that wear resistance was mainly dependent on the grain size and hardness, and a good correlation existed between tool life and hardness for all grades tested. It was also noted in the work that wear of the tungsten carbide tool took place primarily on the clearance face and the amount of wear was proportional to the normal force. The wear of tungsten carbide tipped circular saws was reported in the literature during cutting particleboard [41,42]. It was noted in the work that wear of the cutting edge occurs primarily in the clearance face, with visible grooves developing along the cutting direction. Scanning electron microscopic (SEM) examination of the worn surfaces revealed that wear of the cutting edge occurred by preferential removal of the cobalt binder between the tungsten carbide grains. Such removal caused the tungsten carbide grains on the surface of the tool to be loosely held in the composite matrix and subsequently be mechanically removed from the cutting edge. It was also noted in the work that the width of wear land on the clearance face of the tool was proportional to the normal force component, and suggested that friction on the back surface of the tool was responsible for tool wear in the circular sawing of particleboard [41]. The effects of work materials and grades of tungsten carbide tool on tool wear were investigated during the peripheral milling of particleboard [43]. Six types of melamine coated and uncoated particleboard, and eight types of micrograin tungsten carbide were experimented in the work. It was found that high correlation coefficients existed between tool wear and mineral content of the particleboard. This result implies the significance of mechanical wear by micro-abrasion. Furthermore, increasing the vanadium content of the

10

I. Introduction

tungsten carbide tool was reported to slightly decrease the cutting edge wear, whereas increasing the nickel and chromium content slightly increased tool wear [43]. During machining particleboard and fiberboard, it was observed that wear of tungsten carbide tools proceeds in a similar manner [42]. First, the binder phase was partly removed between the tungsten carbide grains by a combination of plastic deformation and micro-abrasion. The second stage of wear occurs after sufficient binder was removed to allow removal of carbide grains from the surface by mechanical forces. It was also noted in other studies that particleboard and fiberboard often contain abrasive particles, such as dust, sand, silica and cured-resin [43,44,45]. Loose microfragments of these abrasives at the interface between tool and work material was reported to penetrate, under cutting pressure, between the carbide grains and preferentially remove the cobalt binder by extrusion and micro-abrasion. In any given situation, it is likely that more than one mechanism is operative, and that interactions occur between the various mechanisms. However, identification of the wear mechanisms of tungsten carbide tools during cutting wood is not an easy task due to the complexity of the problem, and separating an individual wear mechanism during the cutting process is rather complicated. The influence of mechanical and physical properties of the tungsten carbides and the specific conditions under which the tools are used, such as work materials hardness, moisture content of work materials, and cutting speed, would be expected to cause one mechanism or very few mechanisms to control the rate of wear. It can be considered from the work reviewed above that tool wear takes place by removing the cobalt binder between the tungsten carbide grains, which eventually leads to removal of the carbide grains. Removal of the cobalt binder is basically due to chemical wear mechanism by oxidation or corrosion, and mechanical wear mechanism by micro-abrasion, extrusion or erosion. Corrosion and oxidation involve chemical transformation of tool material into compounds that can be easily removed from the cutting edge by mechanical abrasion. Micro-abrasion, erosion and extrusion involve the mechanical removal of microscopic particles.

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1.5 Improvement in Wear Resistance of Cutting Tools The effects on tool wear of adding corrosion resistance elements to the cobalt binder of the tungsten carbide alloys are reported in the literature [46]. It was noted in the literature that adding elements, such as chromium, nickel and molybdenum, to the binder metal did not bring any significant improvements in wear resistance. In another study [47], boron was added to the cobalt binder of the tungsten carbide by thermal diffusion. It was noted in the work that boriding the tungsten carbide resulted in less tool wear while machining MDF. This improvement was attributed to the increase in hardness and oxidation resistance of the cobalt binder, which was caused by the formation of cobalt borides at the surface of the tool. However, an apparent of disadvantage of boriding treatment by thermal diffusion was reported in which the cobalt borides did not maintain wear resistance for along time. Another effort that has been developed to increase the life of the tools is surface hardening. Some methods of surface hardening were already tried. Hardening the surface of carbide cutting tool with chromium by electric spark deposition method was reported in the literature [48]. Further, plating the surface of carbide tool with chromium was also experimented [49,50]. Other methods are the formation of blue oxide film on the carbide tool surfaces by heating in a steam atmosphere, and treatment of tools in salt baths to introduce high levels of carbon, nitrogen and/or sulphur onto the tool surfaces [24]. However, none of them has become a patent. Further development of coated tools was the application of a thin layer of hard TiC coating to the cemented carbide tool by chemical vapor deposition (CVD) method. The impetus for this development came from the Swiss Watch Research Institute, where vapor-deposited TiC coating had been used on steel watch parts and cases to combat wear on these components. High temperatures employed during CVD coating generally ensure good bonding between the substrate and the coating. However, coating adhesion can be adversely affected by stresses caused by the thermal expansion mismatch between the substrate and the coating. Therefore, in the 1980’s, physical vapor deposition (PVD) was developed as a commercial process for applying hard TiN coatings to cemented carbide tools.

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I. Introduction

Recently, development of these processes has resulted in the commercial availability of cutting tools coated with thin layers of refractory metal carbide or nitride. Coatings at present available consist of titanium carbide (TiC), titanium nitride (TiN), titanium aluminum nitride (TiAlN), titanium carbonitride (TiCN), chromium nitride (CrN), hafnium nitride (HfN), diamond and diamond-like carbon (DLC). These coatings have been widely used in the metal working industry with a good result, however little is known concerning their behavior in wood and wood-based machining. Tungsten carbide inserts were coated with TiN and TiC films by CVD method for cutting particleboard [42]. It was noted in the work that TiN coating brought no advantages in the milling particleboard, i.e. the edge life attained or the total tool paths, respectively were not longer than those attained by uncoated cutting edges. TiC coated tungsten carbide tool was also reported to show in large edge fractures after a short time of operation. Performance of different PVD coated tungsten carbide tools in continuous milling of particleboard was reported in the literature [51]. The tungsten carbides which have different grades were coated with TiN, Ti(N, CN), and TiAlN2. It was noted in the literature that a slight improvement in wear resistance was provided by coatings which were synthesized to carbide grade with fine grain size (0.8 µm) and low cobalt content (3%). In contrast, coatings applied to carbide grade with higher cobalt content and coarse grain size were reported to decrease the wear resistance of the tools. The primary failure mode for the PVD coatings tested in the work was chipping of coatings on the rake face. This fact was caused by inadequate adhesion of coating over the tool surfaces due to improper substrate and coating combination. TiN, TiAlON and TiC coatings were synthesized on the surface of tungsten carbide (93.5% WC, 5% Co, 1.5% (Ta, Nb)C by plasma assisted chemical vapor deposition (PACVD) for milling laminated particleboard [51]. It was noted in the work that TiN and TiAlON coated carbide tools did not provide any improvement in wear resistance compared to uncoated carbide tool, and the TiC coated carbide tool provided only a slight improvement. Performance of CVD diamond coated carbide tools was reported during machining wood and wood-based materials [53,54,55]. It was noted that during cutting plywood, air-dried and wet melapi, delamination did not occur in any of the coated tools, however delamination was observed in all

13


coated tools during cutting particleboard, MDF, and wood-chip cement board [55]. Diamond film was reported to peel off in the early stage of cutting for every feed per revolution performed [53]. It was also noted that the rate of delamination became slower as the film thickness became thicker, and the 15 µm and 20 µm coated tools with co-etching showed excellent durability in comparison with the other coated tools [54].

References 1. Kennedy D, Xue Y, Mihaylova M (2005) Current and Future Applications of Surface Engineering. The Engineers Journal (Technical). Vol. 59: 287– 292. 2. Bunshah RF (2001) Handbook of hard coatings, Noyes Publication. 3. Martin PM (2010) Thin Film Coatings, III-rd Ed., Elsevier Inc. 4. Gregson V (1984) Laser Material Processing, Holland Publishing Company, Holland. 5. Semak V, Dahotre N (1988) Lasers in Surface Engineering, Surface Engineering Series, vol. 1, ed. by N. Dahotre, ASM International, Materials Park, OH, USA 6. Sheehy M, Tull B, Friend C, Mazur E (2007) Chalcogen doping of silicon via intense irradiation, Materials Science and Engineering B. 137(1–3): 289–294 (2007). 7. Lawrence JR, Dowding C, Waugh D, and Griffiths JB (2015) Laser Surface Engineering: Processes and Applications, Woodhead Publishing, Cambridge, USA. 8. Etsion I (2005) State of the Art in Laser Surface Texturing J. Tribol. Trans. ASME 127(1): 248–252. 9. Matthews DTA, Ocel ́ık V, and deHoss JThM (2007) Tribological and mechanical properties of high power laser surface treated metallic glasses, Mater. Sci. Eng. A, Vol. 471: 155–164 10. Jiang W and Molian P (2001) Nanocrystalline TiC powder alloying and glazing of H13 steel using a CO2 laser for improved life of die casting dies, Surface and Coatings Technology, Vol. 135: 139–149.

14

I. Introduction

11. DiMelfi RJ, Sanders PG, Hunter B, Eastman JA, Sawley KJ, Leong KH, and Kramer JM (1998) Mitigation of subsurface crack propagation in railroad rails by laser surface modification. Surface and Coatings Technology. Vol. 106: 30–43. 12. Anonymous (2016) FAO Yearbook of Forest Products. No. 27. Rome. Italy. 13. Robert RA, Cary RA (1980) Tool Steels 4th Edition. ASM International, Materials Park, Ohio. pp 820. 14. Stewart HA (1989) Feasible High-Temperature Phenomena in Tool Wear from Wood Machining. Forest Prod. J. 39(3):25–28. 15. Stewart HA (1992) High-Temperature Halogenation of Tungsten Carbide Cobalt Tool Material when Machining Medium Density Fiberboard. Forest Prod. J. 42(10): 27–31. 16. Stewart HA (1992) A Comparison of High Speed Steels for Wood Machining. Forest Prod. J. 42(7/8): 73–77. 17. Duff KW (1958) Carbide Wood Cutting Tools. Forest Prod. J. 8(5):3336 18. Weill TC (1958) Cemented Tungsten Carbides and their Application to the Wood Working Field. Forest Prod. J. 8(6): 21–24. 19. Reynolds RV (1958) Status of Tungsten Carbide Tools in the Wood Working Industry. Forest Prod. J. 8(5): 24–26. 20. Tanaka C, Takahashi A, Shiota Y (1986) Cutting Performance of Cermet, Ceramic, CBN and Artificial Diamond I (Wear from Continuous Cutting of Wood-Based Materials). Mokuzai Gakkaishi. 32(2): 96–102. 21. Tanaka C, Takahashi A, Date H, Nakao T (1988) Cutting Performance of Cermet, Ceramic, CBN and Artificial Diamond III (Cutting Performance of Ceramic Tools). Mokuzai Gakkaishi. 34(4): 298–304. 22. Eda H, Date H, Tanaka C, Takahashi A, Nakao T (1990) Cutting Performance of Ceramic Tools. Mokuzai Kogyo. 45(5)216–219. 23. Krar SF, Oswald JW, Amand JE (1969). Technology of Machine Tools. Mc Graw-Hill Company. Canada. pp 371–389.

15


24. Trent EM (1996) Metal Cutting. Butterworth-Heinemann. Oxford, pp37–56. 25. Clark IE (1993) PCD as a Tool Material for Wood Working Applications. Proceedings of the 11th International Wood Machining Seminar. pp 96– 113. 26. Inoue H (1988) Edge wear of tools tipped with new materials in the cutting of wood. Mokuzai Gakkaishi (in Japanese). 34:291–297. 27. Szymany R (1997) Wear Resistance Improvement of Cutting Tools Used in Secondary Processing of Incense Cedar. Proceedings of the 13th International Wood Machining Seminar. pp 641–650. 28. Kirbach E, Chow S (1976) Chemical Wear of Tungsten Carbide Cutting Tools by Western Red Cedar. Forest Prod. J. 26(3):44–48 29. Murase Y (1984) Effect of Tool Materials on the Corrosive Wear of Wood-Cutting Tools. Mokuzai Gakkaishi (in Japanese). 30: 47–54. 30. Murase Y (1986) Effect of Tool Materials on the Corrosive Wear of Wood-Cutting Tools II. Mokuzai Gakkaishi (in Japanese). 32: 596– 602. 31. Morita T, Banshoya K, Tsutsumoto T, Murase Y (1999) Corrosive Wear Characteristics of Diamond-Coated Cemented Carbide Tools. J. Wood Science. 45: 456–460. 32. Kirbach E, Bonac T (1982) Dulling of Saw Teeth Tipped with a Satellite and Two Cobalt-Cemented Tungsten carbides. Forest Prod. J. 32(9): 42– 45. 33. Fukuda H, Banshoya K, Murase Y (1992) Corrosive wear of WoodCutting Tools I. Effect of Tool Materials on the Corrosive Wear of Spur Machine Bits. Mokuzai Gakkaishi. 38: 764–770. 34. Fukuda H, Banshoya K, Manatani T, Murase Y (1994) Corrosive wear of Wood-Cutting Tools II. Effect of Alloys Compositions on the Corrosive Wear of Cemented Carbide Bits. Mokuzai Gakkaishi. 40: 687–693. 35. Padilla MH, Rapp RA, Stewart HA (1991) High-Temperature Oxidation of Tungsten Carbide-Cobalt Composites in the Presence of MDF. Forest Prod. J. 41(10): 31–34.

16

I. Introduction

36. Stewart HA, Shatynski SK, Harbison B, Rabin B (1986) HighTemperature Corrosion of Tungsten Carbide from Machining Medium Density Fiberboard. The carbide and Tool. 18(1): 2–7. 37. Kim MG, Stewart HA, Wan H (1999) Effects of Anionic Components of Urea Formaldehyde Resins Used as Particleboard Binders. Forest Prod. J. 49(4): 60–65. 38. Sheikh-Ahmad JY, Bailey JA (1999) On the Wear of Cemented Carbide Tools in the Continuous and Interrupted Cutting of Particleboard. Proceedings of the 14th International Wood Machining Seminar. pp 211– 221. 39. Sheikh-Ahmad JY, Bailey JA (1999) The wear characteristics of some cemented tungsten carbides in machining particleboard. Wear (1999). 256–266. 40. Salje E, Stuehmeier W (1988) Milling Particleboard with High Hard Cutting Materials. Proceedings of the 9th International Wood Machining Seminar. pp 211–228. 41. Sugihara H, Okumura S, Haoka M, Ohi T, Makino Y (1979) Wear of Tungsten Carbide Tipped Circular Saws in Cutting Particleboard: Effect of Carbide Grain Sizes on wear Characteristics. Wood Science and Technology. 13: 283–299. 42. Sheikh-Ahmad JY, Bailey JA (1999) High-Temperature Wear of Cemented Tungsten Carbide Tools while Machining Particleboard and Fiberboard. J. Wood. Science. 45: 445–455. 43. Porankiewicz B (1997) Variation in Composition of Micrograin Cemented Carbide and its Impact on Cutting Edge Wear during Particleboard Machining. Proceedings of the 13th International Wood Machining Seminar. pp 791–799. 44. Porankiewicz B, Gronlund A (1991) Tool Wear Influencing Factors. Proceedings of the 10th International Wood Machining Seminar. pp 220–229. 45. Huber H (1985) Tool Wear Influenced by the Contents of Particleboard. Proceedings of the 8th International Wood Machining Seminar. pp 72– 85.

17


46. Reid A, Stewart HA, Rapp A (1991) High-Temperature Reactions of Tungsten Carbide Cobalt Tool Material with MDF. Forest Prod. J. 41(11/12):12–18. 47. Stewart HA (1987). Borided Tungsten Carbides Reduces Tool Wear during Machining of MDF. Forest Prod. J. 37(7/8):35–38. 48. Halvorson HN, Stuart WMP (1963) Improvement of sawmill cutting tool sharp life by surface hardening. Forest Prod. J. 37(13): 108–110. 49. Kato C, Kawai Y, Soga K, Fukui H (1990) The wear characteristics of a wood working knife with chromium plating I: The effect of coating condition in the orthogonal cutting of transverse surfaces of wood. Mokuzai Gakkaishi (in Japanese), 36: 615–623. 50. Kato C, Kawai Y, Soga K, Fukui H (1990) The wear characteristics of a wood working knife with chromium plating II: The influence of depth of cut on the self-sharpening characteristics in orthogonal cutting of transverse surfaces of wood. Mokuzai Gakkaishi, 36: 844–850. 51. Sheikh-Ahmad JY, Stewart JS (1995) Performance of Different PVD Coated Tungsten Carbide Tools in the Continuous Machining of Particleboard. Proceedings of the 12th International Wood Machining Seminar. pp 282–291. 52. Fuch I, Raatz Ch (1997) Study of wear behavior of specially coated (CVD, PACVD) cemented carbide tools while milling of wood-based materials. Proceedings of the 13th International Wood Machining Seminar. pp 709–715. 53. Morita T, Banshoya K, Tsutsumoto T, Murase Y (1995) Cutting Performance of Diamond-Coated Cemented Carbide Tools. Proceedings of the 12th International Wood Machining seminar. pp 1093–1101. 54. Morita T, Banshoya K, Tsutsumoto T, Kawamitsu M, Murase Y (1997) Characteristics of Diamond-Coated Cemented Carbide Tools in the Milling of Particleboard. Proceedings of the 13th International Wood Machining seminar. pp 1093–1101. 55. Morita T, Banshoya K, Tsutsumoto T, Murase Y (1998) Effect of Work Materials on Cutting Performance of Diamond-Coated Cemented Carbide Tools. Forest Prod. J. 48(5):43–50.

18

II. The Importance of Extractives and Abrasives in Wood Materials on the Wearing of Cutting Tools High-speed steel and tungsten carbide cutting tools are widely used in the woodworking industry throughout Indonesia for machining tropical woods and wood composites, which are mostly high in extractives and abrasive materials. A few studies that extensively explore chemical components of some of the commercial wood species have been done. It was noted that their extractives range from 1.4% to 13.8%, and their ash-contents range from 0.1% to 5.0% [1]. It is known that a chemical component of woods may play an important role in determining cutting tool wear rates. To adequately describe the wear mechanisms, wood and cutting tool interactions must be considered. This is not a simple problem due to the large number of possible interactions, as well as the difficulties in characterizing each of these interactions. There is evidence that the extractives in wood act as lubricants and effectively decrease the coefficient of friction during wood cutting [2]; however, many studies indicate that the extractives in wood cause adverse effect on wearing of cutting tools. Hillis and McKenzie [3] postulate a chelation reaction between polyphenolic extractives and the iron in steel cutting tools as a wear mechanism. McKenzie and Hillis [4] produced etching on steel knives by exposing them to chemical solutions typical of those found in wood. Their results showed that a measurable amount of tool wear was due to chemical reaction between the cutter and the wood being cut. Kirbach and Chow [5] emphasized the complexity of tool wear problem inherent in the multi-component nature of the work and tool materials. It was noted in their study that the wear of carbide tool was due to a chemical attack of the tool material binder and a mechanical failure of the exposed carbide grains.


At the period after 1980, chemical wear due to extractives in the woods, such as gums, fats, resins, sugars, oils, starches, alkaloids and tannins, have also been reported as an important factor in determining the chemical wear of woodworking cutting tools [6,7,8,9,10]. Extractives vary in chemical composition, amount, and reactivity among tropical wood species which might affect their degree of chemical impact on cutting tool materials. Rapid mechanical wearing of cutting tools has often been attributed to the presence of silica and other abrasive agents in the woods [11,12,13,14]. It was noted in these studies that woods with high silica content caused high wear rate of high-speed steel cutting tool. However, the authors did not explain in detail the importance of silica content and its distribution on the wearing phenomenon of the high-speed steel cutting tool. The Indonesian industry is now about to utilize lesser known species and fast growing wood species for construction, furniture, and wood composite products (wood cement board, particleboard, medium density fibreboard (MDF), and oriented strand board (OSB)). These wood composites are made of mix fast growing wood species. The focus of this chapter is to describe the effect of extractive and silica contained in a lesser known species and wood composites on chemical and mechanical wear characteristics of high-speed steel and tungsten carbide cutting tools.

2.1 Chemical Wear Caused by Extractives Extractive and silica content in the wood materials, as well as the percentage of weight loss of tool materials after 48 hours of exposure to wood material extracts at 80oC are presented in Table 2.1 and Table 2.2, respectively. The amount of chemical wear in this work was determined by the percentage of weight loss of the tool materials. The results in Table 2.1 show that the TapiTapi wood was acidic and the wood composites tested were nearly neutral in pH. The wood materials varied slightly in extractive content. Tapi-Tapi wood had the highest extractive content, and wood cement board had the highest silica content. The results in Table 2.2 indicate that SKH51 and K10 tool materials suffered weight losses for all wood materials. High extractive content of these wood materials could be the reason for the weight loss of the tool materials [15].

20

II. The Importance of Extractives and Abrasives in Wood Materials on the Wearing of Cutting Tools

High extractive and acidity of Western red cedar [5] and of eucalypt [7] were noted to cause corrosion of steel cutting tools. The SKH51 and K10 tool materials suffered the highest percentage of weight loss when they were soaked in Tapi-Tapi wood extract. Though the amount of extractive per unit oven dry volume in the Tapi-Tapi wood is almost the same with the others, the extractives of Tapi-Tapi wood had the strongest acidity, however. This indicates that the strong acidity of chemical compounds in the Tapi-Tapi wood extract compared to the others is more reactive to elements of the tool materials [15]. It was investigated that the surface of the tip of SKH51 after soaked in Tapi-Tapi extract was covered by a light brown compound. Under SEM/EDS analysis, the light brown compound revealed some chemical elements dominated by iron oxide (Fe and O) as shown in Figure 2.1a. The presences of Fe and O indicated the occurrence of corrosion on the surface of the tip of tool materials. The corrosion could take place on certain parts of the surface of the tip as indicated by high peaks of O profile in Figure 2.1b. Table 2.1 Chemical Characteristics of the Experimented Wood Materials Wood materials Tapi-tapi Wood cement board Particleboard MDF OSB

pH 4.47 7.78 6.61 7.05 6.16

Extractive (%) 13.8 9.6 11.6 10.9 8.8

Ash (%) 1.27 94.2 2.87 6.0 2.41

Silica (%) 0.75 55.1 1.20 0.95 0.47

Table 2.2 Percentage of Weight Loss of Tool Materials after 48 Hours Reaction at 80°C with Wood Materials Extract Tool materials SKH51 K10

Wood extract Tapi-Tapi wood Wood cement board 0.48 0.11

0.06 0.01

Particleboard MDF OSB 0.08 0.02

0.11 0.02

0.14 0.05

21


cps

O profile

Fe

200 150 100 O 50 0

W Mo 2

V C Mn r 4 6

(a)

Fe

Fe W 8

10

Energy (keV)

120 µm

(b)

Figure 2.1 SEM/EDS analysis indicating corrosion on the surface of the SKH51 tool material after reaction with the Tapi-Tapi wood extract Table 2.3 Specification of Cutting Tool Materials Tool materials

Dimension

SKH51 high-speed steel

2x10x20 mm

K10 Tungsten carbide

2x10x20 mm

Specification Metal components Heat treatment Hardness C=0.88, Si=0.25, Hardened at Mn=0.30, P=0.02, 1220 oC followed 815 HV0.5 S=0.001Cr=4.04, by two times W=6.13, Mo=4.92, of one hour V=1.85%wt tempering at 560oC with cooling in air WC=94%, Co=6%wt 1450 HV0.5

The SKH51 tool materials suffered higher weight loses when compared to K10 tungsten carbide for all wood materials. This phenomenon is considered to be the result of wide variation in metal components of the SKH51 [14]. The iron (Fe) in the SKH51 was susceptible to suffer corrosion as attacked by reactive chemical compounds of wood extractive, rather than tungsten carbide (WC) and cobalt (Co) in the K10 tungsten carbide (Table 2.3). This fact could also be related to the phenomenon that the microstructure of the tungsten carbide tools, which consists mainly of austenitic and martensitic matrix, are more stable and more resistance to chemical reaction than that of the hardened steel tools, which consists mainly of ferrite [16].

22


2.2 Mechanical Wear of Tool Bits The amount of mechanical wear was determined by the amount of knife-edge recession and abrasion on the clearance face of the bits. Wear behaviour on the clearance face of the bits is presented in Figure 2.2, and wear pattern of the bits is presented in Figure 2.3. The wear rates of the bits obtained from the linear regression equation in Figure 2.2 are summarized in Table 2.4.

Figure 2.2 Wear behaviors of the SKH51 and K10 bits with cutting length in routing wood materials The results in Figure 2.2 depict that wood cement board and particleboard wore the SKH51 and K10 bits faster when compared to the other wood materials. Wood cement board and particleboard also caused the largest rate of wear of the bits tested (Table 2.4). High-speed steel suffered a remarkable fracture edge wear at 300 m cutting length in the wood cement board. The cutting test of the high-speed steel for wood cement board was stoped at the 300 m cutting length due to the serious damage of the cutting edge. This is due to the much higher content of silica, which has an average hardness of about 1200 HV, in the wood cement board compared to that in the other wood materials tested (Table 2.1).

23


Table 2.4 Rate of Wear (µm/km) of the Router Bits for Different Wood Materials According to Linear Function in Figure 2.2 Bits SKH51 K10

Tapi-tapi 53.9 17.6

Wood materials Cement board Particleboard 150,0 58.4 27.9 27.3

MDF 54.4 20.9

OSB 39.6 17.1

It was visually investigated that the hard particles of cured cement distribute evenly in the matrix of the wood cement board. The cement used in the production of this wood cement board consisted of 75% limestone (CaCO3), 20% silica (SiO2) and alumina (Al2O3), and 5% incidental ingredients (SO3 and MgO). The SEM photomicrographs of particleboard and OSB also revealed cured thermosetting resin. It was investigated that the outer and inner parts of particleboard contained higher dosage of the thermosetting resin of urea-formaldehyde when compared to OSB and MDF. The cured cement and thermosetting resin were considered to impose severe mechanical abrasion on the cutting tool edges during the cutting the wood cement board and particleboard. The SEM micrograph of Tapi-Tapi wood in Figure 2.4a reveales a few round crystals. EDS analysis indicated the round crystals consisted of Si and O (Figure 2.4a right). This analysis confirms the previous result that silica compounds in wood of the form of silicon dioxide (SiO2) [17]. In addition, silicon dioxide in woods occurred at the inter-layer of tracheids, ray parenchymas, and bordered pits without any geometrical form [18]. It appears under high magnification that the surface of the silica crystals in the Tapi-Tapi wood was corrugated, which caused high mechanical abrasion when cutting. The results in Figure 2.4b-e (right) show EDS analysis of elemental profiling on randomly selected spectra at the surface of the SEM micrographs. The presence of Si and O in the spectra indicated the presence of silica compound in the tested wood materials. EDS analysis on the SEM micrograph in Figure 6b shows that the cured cement consisted of Ca, Si, O, Al, K, Mg, C. This EDS result confirmed the essential raw ingredients of cement (limestone, silica, and alumina) [15]. The EDS results from the SEM micrographs of particleboard, MDF, and OSB also revealed the presence of Si and O, which indicated the presence of silica.

24


(a)

(b)

(c)

(d)

(e)

0 km

2 km SKH51

2 km

0 km K10

Figure 2.3 Wear patterns of the SKH51 high-speed steel and K10 tungsten before and after 2 km cutting length for Tapi-Tapi wood (a), wood cement board (b), particleboard (c), MDF (d), and OSB (e)

25


cps

EDS

Si

100 80

C

60 40

a

20

O Energy

0

2

cps

4

6

10

C

100

EDS

8

80 60 40

b

20

O C

Mg Al

0

Si

2

cps

4

Energy

6

8

10

Si

100

EDS

C

Cl

80 60

O

40

c

20 0

C

N

M

Al

2

cps

4

100

EDS

40

d

6

Ca

Energy 8

10

Si

80 60

K Ca

Ca C

O

20

K

0

2

cps

4

6

Energy 8

10

100 80 60 40

e

20 0

CO

Ca Si Mg Al 2

P

Ca

S 4

6

M Fe Energy 8 10

Figure 2.4 SEM/EDS analysis of the wood materials showing the abrasives for Tapi-Tapi wood (a), wood cement board (b), particleboard (c), MDF (d) and OSB (e)

26


The occurrences of silica in Tapi-Tapi wood, wood cement board, particleboard, MDF, and OSB at every 500 µm2 spectra on the SEM photomicrographs gives an indication that the cutting edge of the bits would engage the silica during cutting these wood materials [15]. The higher frequency of occurrence of silica in the wood cement board, particleboard and Tapi-Tapi wood would cause the cutting edge of the bits to be dulled at a faster rate compared to MDF and OSB.

Rate of Wear (µm/km)

90,0

SKH51 K10

75,0

y = 24,4x + 29,4 r = 0,96

60,0 45,0

y = 11,9x + 9,4 r = 0,66

30,0 15,0 0,0

0,4

0,6

0,8

1,0

1,2

1,4

Silica Content (%) Figure 2.5 Relationship between rate of clearance wear and silica content for SKH51 and K10 bits. Note: y (Rate of Wear), x (silica content), r (correlation coefficient) It appears from the results in Figure 2.2 and Table 2.4 that the clearance wear and the wear rate of the SKH51 bit were twice as large compared to those of the K10 tungsten carbide bit when cutting the same wood materials. Further, the results in Figure 2.5 give an indication that the rate of clearance wear for both SKH51 and K10 increased linearly with increasing in the silica content. However, the progress of wear rate with increasing in the silica content was quite different. The clearance wear of the SKH51 would increase in faster rate compared to that of the K10 carbide with increasing in the silica content of the work materials. The lower resistance to chemical attack by extractive and the lower in hardness of the SKH51 compared to that of K10 carbide bit (Table 2.1) is considered as the reason for this phenomenon.

27


SEM spectra

worn cutting edge

SKH51

K10

Figure 2.6 SEM micrograph of the worn cutting edge of the high-speed steel (upper) and tungsten (under) in cutting Tapi-Tapi wood at the 2 km cutting length The results in Figure 2.3 give an indication that wear patterns of the SKH51 bits were the same for all wood materials, except for wood cement board. The cutting edge with sharpness angle of 65o used in this experiment was not strong enough for the SKH51 bit to cut the wood cement board [15]. The initial failure of the cutting edge occurred due to severe fracture wear. Therefore, it could be recommended to use a sharpness angle around 90o with a negative rake angle, especially for the cutting the wood cement board. Considering the results in the previous paper [19] and the facts that the wood materials tested are air-dried and are routed at a low cutting speed, the wear patterns of the bits in Figure 2.3 were predominantly caused by mechanical abrasion. However, there was a slight difference in clearance wear pattern between the K10 tungsten carbide and SKH51 high-speed steel. The cutting edge of K10 tungsten carbide compared to SKH51 was more corrugated, especially when cutting the wood cement board and particleboard (Figure 2.6

28


under). SEM micrograph of the K10 revealed that the corrugated edge was caused by lower in toughness of the K10 tungsten carbide compared to the SKH51, which leads to retraction of carbide grain from the fraction.

2.3 Importance of Silica and Extractive in Wearing of Cutting Tools Mechanical and chemical interactions of silica and extractives onto cutting tool surface played an important role in the process of wearing of SKH51 and K10. The wearing of the SKH51 and K10 cutting edges was observed due to corrosion and abrasion mechanisms. The dominant type of wear in routing the Tapi-Tapi, wood cement board, particleboard, MDF, and OSB was mechanical abrasion due to silica. However, corrosion/oxidation mechanism was also found to provide an important contribution to the wear of SKH51 and K10 when cutting the Tapi-Tapi, wood cement board, particleboard, MDF, and OSB. The results in Figure 2.7 indicate that the silica and extractive determined the wear rate of the SKH51 and K10 to a varying degree. Though Tapi-Tapi wood compared to MDF is lower in density and silica content, they caused almost the same rate of wear to the SKH51, however. This phenomenon gives an indication that corrosion mechanism was to be a major contributor to the wear of SKH51 when cutting the Tapi-Tapi wood. In this study, high percentages of Fe and O were detected in the surface of the SKH51 after soaked in the extract of Tapi-Tapi wood (Table 2.2, Figure 2.1). Further, the Fe and O were observed to form red brown compounds of iron oxide on the surface of cutting edge of SKH51 after routing the Tapi-Tapi wood [15]. The red brown compounds were easily removed by mechanical abrasion as the cutting edge of the SKH51 engaged the Tapi-Tapi wood. Therefore, the presence of corrosion will promote faster rate of cutting tool wear. However, when the Tapi-Tapi was routed using K10 cutting tool, corrosion /oxidation mechanism was to be a minor contributor to the wear of the K10. It also appears from the results in Figure 2.7 that silica content of 0.47% in OSB and of 0.75% in Tapi-Tapi wood caused almost the same wear rate to the K10 cutting tool. This result indicates that wear mechanisms other than abrasion by silica and corrosion by extractive may contribute to the wear of K10 when cutting the OSB. In this study, OSB and particleboard machined consisted of urea formaldehyde resin, additives, and catalys.

29


70 60 50 40 30 20 10 0

SKH51

8.8%

13.9%

0.47

0.75

10.9%

11.6%

0.95

1.20

Silica content (%)

Wear Rate (µm/km)

Wear Rate (µm/km)

Kim et. al. [20] stated that the acidity effect and corrosion/oxidation mechanisms of anionic components of additives and catalys are responsible for an increase in cutting tool forces and cutting tool wear. 70 60 50 40 30 20 10 0

K10

8.8%

13.9%

10.9%

11.6%

0.47

0.75

0.95

1.20

Silica content (%)

Figure 2.7 Interaction effect of silica and extractive on the wearing rate of SKH51 and K10 cutting tools. Note: extractive content of 8.8% (OSB), 10.9% (MDF), 11.6% (particleboard), 13.9 (Tapi-Tapi) Abrasion has been considered to be the major tool wear mechanism in cutting the MDF, particleboard and wood cement board in this study. However, machining wood based products such as MDF, particleboard and wood cement board also involve high temperatures and presures near the cutting tool edge. In the previous study, temperatures were measured and estimated to be 300–400oC at the cutting tool edge when cutting MDF and wood cement board at 20 m/s cutting speed [19]. Further, Reid et. al. [21] noted that thermal decomposition of MDF was evident at temperature of 325oC. The results of energy dispersive spectroscopy (EDS) mapping of worn cutting tool edge in these studies indicated that oxidation/corrosion mechanism contributed significantly to the wear of tungsten carbibe cutting tools when cutting MDF and wood cement board.

References 1. Martawijya A, Kartasujana I, Kadir K, and Prawira S (1989). Atlas Kayu Indonesia, Forest Products Research Institute, Bogor. 2. McKenzie WM, and Karpovich H (1968) The frictional behaviour of wood, Wood Science and Technology 2(2), 139–152. 30


3. Hillis WE, and McKenzie W M (1964) Chemical attack as a factor in the wear of woodworking cutters, Forest Products Journal. 14(7) 310–312. 4. Mackenzie WM, and Hillis WE (1965) Evidence of chemical acceleration of wear in cutting plant materials, Wear 8(3), 238–243.

5. Kirbach E, and Chow S (1976) Chemical wear of tungsten carbide cutting tools by western red cedar, Forest Products Journal. 26(3): 44– 48. 6. Fukuda H, Banshoya K, and Murase Y (1992) Corrosive wear of woodcutting tools I: Effects of tool materials on the corrosive wear of spur machine bits, Mokuzai Gakkaishi. 38(8): 764–770. 7. Kirlov A (1986) Corrosion and wear of sawblade steels, Wood Science and Technology 20(4). 361–368. 8. Morita T, Banshoya K, Tsutsumoto T, and Murase Y (1999) Corrosive wear characteristics of diamond-coated cemented carbide tools, Journal of Wood Science. 45(6): 456-–460. 9. Murase, Y. (1984). “Effect of tool materials on the corrosive wear of wood-cutting tools,” Mokuzai Gakkaishi 30(1): 47–54. 10. Darmawan W, and Tanaka C (2006) Chemical and mechanical wearing of woodworking cutting tools by tropical woods. Journal of Tropical Forest Science. 18(4): 166–172. 11. Hayashi K, Suzuki T (1983) Effect of cutting speed on tool wear in the peripheral milling of wood (in Japanese). Mokuzai Gakkaishi. 29(1): 36– 42. 12. Huber H (1985) Tool wear influenced by the contents of particleboard, In: Proceedings of the 8th International Wood Machining Seminar. 7–9 October 1985, California, pp. 72–85. 13. Porankiewicz B, and Gronlund A (1991) Tool wear-influencing factors, In: Proceedings of the 10th International Wood Machining Seminar. 2–4 October 1991, Kyoto, pp. 220–229 14. Darmawan W, Rahayu I, Nandika D, Marchal R (2011) Wear Characteristics Of Wood Cutting Tools Caused By Extractive And Abrasive Material In Some Tropical Woods. Journal of Tropical Forest Science. Vol 23 (3): 345–353.

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15. Darmawan W, Rahayu I, Nandika D, Marchal R (2012) The Importance of Extractives and Abrasives in Wood Materials on the Wearing of Wood Cutting Tools. Journal of BioResources. Vol 7(4): 4715–4729. 16. Pipple E, Woltersdorf J, Pockl G, and Lichtenegger G (1999) Microstructure and nanochemistry of carbide precipitates in high-speed steel S6-5-2-5, Materials Characterization. 43(1): 41–55. 17. Misra MK, Ragland KW, and Baker AJ (1993) Wood ash composition as a function of furnace temperature, Biomass and Bioenergy .4(2): 103– 116. 18. Xiao-mei J and Yin Z (1989) SEM observation on crystals and silica in wood species of Chinese Gymnospermae. Acta Botanica Sinica. 31(11): 835–840. 19. Darmawan W, Tanaka C, Usuki H, and Ohtani T (2001) Performance of coated carbide tool in turning wood-based material: Effect of cutting speeds and coating materials on the wear characteristics of coated carbide tools in turning wood-chip cement board. Journal of Wood Science. 47(5): 342–349. 20. Kim MG, Stewart HA, and Wan H (1999) Effects of Anionic Components of Urea Formaldehyde Resins Used as Particleboard Binders, Forest Products Journal 49(4):60-65 21. Reid A, Stewart HA, and Rapp A (1991) High-Temperature Reactions of Tungsten Carbide Cobalt Tool Material with MDF. Forest Products Journal. 41(11/12):12–18. 22. TAPPI (1991a) TAPPI Test Methods: Ash in Wood and Pulp (T211 om-85), Volume 1. TAPPI Press, Atlanta. 23. TAPPI (1991b) TAPPI Test Methods: Solvent Extractives of Wood and Pulp (T 204 om-88), Volume 1. TAPPI Press, Atlanta.

32

III. Machining Characteristics of Monolayer Coated Tools in Cutting Hardboard 3.1 Wear and Delamination Characteristics High density fiberboard (hardboard) is generally machined before placed in service. In woodworking industry, cutting tools used in machining woodbased materials are widely cemented carbide tools. However, the use of the cemented carbide tools in some applications involving particleboard and fiberboard is limited because of relatively high rate of wear. Coated carbide and P30 carbide tools have been tested for cutting high density hardboard (Table 3.1). The wear behaviors of rake and clearance faces of the P30 carbide tool and some of the coated carbide tools are provided in Figure 3.1 and Figure 3.2, respectively. Table 3.1 Specifications of coated carbide tools tested, work material machined and cutting conditions Coating material CrC TiC C-TiCN

Deposition temp. (oC) 1000 1000 1000

PVD

TiN P-TiCN VN CrN TiAlN

Work material

Thickness (mm)

Density (g/cm3 )

Hardboard

25

1.18

500 500 500 500 500 Moisture Content (%) 7.4

Cutting tool P30 carbide Coated carbide tools

Coating method CVD

Hardness (HV) 1450 1600 3800 2600

Film thickness (µm) 3–4 3–4 3–4

2000 3–4 3000 3–4 2000 3–4 1800 3–4 2700 3–4 Compressive Shear Hardness strength strength (N/mm2) 2 2 (N/mm ) (N/mm ) 118.8 4.7 70.9


Table 3.1 Specifications of coated carbide tools tested, work material machined and cutting conditions (continue) Cutting condition

Cutting speed (m/min)

Feed speed (mm/rev)

Width of cut (mm)

Geometry

1000

0.1

5.2

Rake angle = 10o, Side rake angle = 2.5o, Clearance angle = 8o

Rake wear (Pm)

The results in Figure 3.1 and 3.2 indicate that the amount of rake and clearance wear increased with increasing in cutting length. The coated carbide tools provided better performance especially in reducing the progression of edge wears than the P30 carbide tool in cutting the hardboard [1,2]. The P30 carbide tool attained rake wear of about 55 µm in cutting hardboard at the 2 km cutting length.Though the coated carbide tools showed almost same rake wears progress, however their amounts of rake wear were slightly different. It appears that the rake wear of the TiC, and TiCN coated carbide tools by CVD method (C-TiCN) were slightly higher than that of the other coated carbide tools by PVD method (TiN, VN, P-TiCN), as shown in Figure 3.1. TiC coated carbide tool attained rake wear of about 52 µm at the 2 km cutting length. 60

P 30 carbide

VN

TiC

TiN

50

P-TiCN

C-TiCN

40 30 20 10 0.0

0.5

1.0

1.5

2.0

Cutting length (km) Figure 3.1 Rake wear progress curves of the P30 carbide tool and some of the coated carbide tools

34

III. Machining Characteristics of Monolayer Coated Tools in Cutting Hardboard

It is observed from the results in Figure 3.2 that the coated carbide tools were also the same in clearance wear progress during cutting the hardboard. The clearance wear of the P30 carbide tool was about 101µm for cutting the hardboard at the 2 km cutting length. The clearance wears at the 2 km cutting length varied from 52 µm to 65 µm for the coated carbide tools, as shown in Figure 3.2. It also appears that all tools showed almost the same wear progress near beginning of the cutting. However, the clearance wear of the P30 carbide tool increased markedly and exceeded the clearance wear of the coated carbide tools, which retained gradual wear progresses during cutting the hardboard, after 1 km cutting length. This fact is probably due to the P30 carbide tool being lower in hardness compared to that of the coated carbide tools. This resulted in the P30 carbide tool having low resistance to the mechanical abrasion than the coated carbide tools investigated.

Clearance wear (Pm)

110

Uncoated Tungsten CrN

90

TiCN

70

TiAlN

TiN

50 30 10 0.0

0.5

1.0

1.5

2.0

Cutting length (km) Figure 3.2 Clearance wear progress curves of the P30 carbide tool and some of the coated carbide tools Figure 3.3 and 3.4 give an indication that the amount of both rake and clearance wears of the coated carbide tools, respectively was closely related to the extent of their delamination during cutting 2 km of the hardboard. These Figures show that almost the same behavior occurred on the coated carbide tools. The amounts of both rake and clearance wear of the coated carbide tools

35


increased with increasing in the delamination on their rake and clearance faces, respectively [1,2]. It appears that the TiC coated carbide tool, which attained slightly higher rake wear among the coated carbide tools investigated, suffered slightly larger delamination of coating film on the rake face during cutting the hardboard (Figure 3.3). When delamination of the TiC coated carbide tool was about 84 µm, its rake wear was about 52 µm. It is observed from the results in Figure 3.1 and 3.2 that the coated carbide tools by PVD method suffered slightly less rake and clearance wear, respectively than the coated carbide tools by CVD method during cutting 2 km of the hardboard. Further, the results in Figure 3.3 and 3.4 give an indication that the coated carbide tools by PVD method also tended to suffer less delamination on both rake and clearance faces, respectively compared to the coated carbide tools by CVD method during cutting the hardboard.

Rake Wear (Pm)

60

TiN VN TiC P-TiCN C-TiCN

50 40 30 20 10 10

30

50

70

90

Delamination (Pm) Figure 3.3 Rake wear and delamination relationship for some of the coated carbide tools As the TiCN coated carbide tool by PVD method (P-TiCN) was compared to the C-TiCN coated carbide tool, the P-TiCN coated carbide tool was lower than the C-TiCN coated carbide tool in both the amount of wear (Figure 3.1, 3.2) and delamination (Figure 3.3, 3.4) during cutting the hardboard [1,2]. This behavior is probably due to the differences in hardness and in the coating

36


process between them. The lower in hardness of the C-TiCN coated carbide tool, which is considered to be caused by its deposition temperature being higher than that of the P-TiCN coated carbide tool (Table 2.2.1), resulted in lower resistant to abrasion. This high temperature was also reported to cause more degradation of the carbide substrate and formation of brittle phases in the interface layer [3, 4, 5, 6], which decrease the tool’s toughness and cause detrimental effects. Therefore, selecting a proper coating process is important in minimizing delamination in order to prolong the tool life.

Clearance Wear (Pm)

70

TiN VN TiC P-TiCN C-TiCN

60 50 40 30 20 10 10

30

50

70

90

110

Delamination (Pm) Figure 3.4 Clearance wear and delamination relationship for some of the coated carbide tools

3.2 Effect of Tool Material Hardness on the Wear and Delamination Figure 3.5 shows that the rake and clearance wear of the tools investigated decreased with an increase in the hardness of the coating materials up to about 3000 HV and then the wears increased slightly when the tool hardness increased in cutting the hardboard at the 2 km cutting length. Decreasing in the amount of wears with increasing the hardness of tool materials from about 2000 HV–3000 HV was not remarkable. The coated carbide tools showed the progression of delamination in the same manner as the progression of their wears. 37


It is observed from the results in Figure 3.5 that the CrC coated carbide tool, which has the lowest hardness among the coated carbide tools, exhibited the highest amount of both rake and clearance wears. Though the TiC coated carbide tool is the largest in hardness among them, its rake and clearance wears was almost the same as those of the CrC coated carbide tool. This fact is attributed to that the cutting edge of the TiC coated carbide tool is probably brittle due to the effect of high temperature during the coating process and is less resistance to high temperature oxidation (Table 3.1). This may result in delamination of TiC coating film to occur more severely than that of the other coated tools investigated, as shown in Figure 2.3. It was noted in another study that the TiC coated carbide tool exhibited premature tool failures due to excessive chipping and insufficient edge strength [5]. These problems were traced to the formation of britle etaphase at the coating/ substrate interface which was caused by surface decarburization during the TiC deposition process [5]. It can be considered from the results in Figure 3.1 through Figure 3.5 that the P-TiCN coated carbide tool would be best in wear and delamination resistance among other tools investigated.

Wear and Delamination (Pm)

120 100 80 60 40 20 0 1000

Rake wear Clearance wear Delamination on rake face Delamination on clearance face

2000

3000

4000

Hardness of Tool Material (Hv) Figure 3.5 Relationship between wears, delamination and hardness of the tool materials at 2 km cutting length

38


3.3 Force and Noise Characteristics of the Coated Tools The results in Figure 3.6 and 3.7 give an indication that though a considerable difference in normal force and noise level, respectively among the coated carbide tools investigated was not observed during cutting the hardboard, however, the normal force and the noise level generated by the coated carbide tools during cutting the hardboard were lower than those of the P30 carbide tool when cutting length exceeded 0.8 km and 0.2 km for normal force and noise level, respectively. The normal force and the noise level generated by the tools increased in proportion with an increase in cutting length and the P30 carbide tool showed the largest proportion especially near end of cutting (Figure 3.6 and 3.7). Therefore, the coated carbide tools are considered to be more advantageous not only in the wear resistance, but also in the noise level and normal force than P30 carbide tool with increasing in cutting length for the high density hardboard.

Normal Force (N)

60 50 40 P30 Carbide P-TiCN C-TiCN TiC

30 20 10 0 0.0

0.5

1.0

1.5

2.0

Cutting Length (km) Figure 3.6 Normal force behaviors with increasing cutting length for P30 carbide and some of the coated carbide tools

39


Noise Level (dB-C)

112 110 108 106

P30 Carbide P-TiCN C-TiCN TiC

104 102 100 0.0

0.5

1.0

1.5

2.0

Cutting Length (km) Figure 3.7 Noise level behaviors with increasing cutting length for P30 carbide and some of the coated carbide tools The behaviors of normal force and noise level of the coated carbide tools (in average) and the normal force and noise level of the P30 carbide tool with clearance wear in cutting the hardboard are shown in Figure 3.8 and 3.9, respectively. The results in these Figures give indication that the normal force and noise level of both the P30 carbide and the coated carbide tools increased linearly with increasing in the clearance wear [7]. However, their progresses in the normal force and noise level were quite different. The normal force (Figure 3.8) and noise level (Figure 3.9) of the coated carbide tools (in average) increased in faster rate compared to those of the P30 carbide tool with increasing in the clearance wear in cutting the hardboard. This fact is probably caused by the uneven surfaces of the cutting edge of the coated carbide tools due to ununiform delamination of the coating films along the cutting edges. Figure 3.10 gives an indication that the normal force was closely related to the noise level during cutting the hardboard with high coefficient of correlation. Normal force of the tools investigated (in average) increased linearly with an increase in their noise levels (in average). When the noise level increased in 3 dB, the normal force of the tools investigated would increase about 11 N.

40


Normal Force (N)

60 y = 0.29x + 21.8 r = 0.92

50 y = 0.41x + 18.2 r = 0.99

40 P30 carbide

30

Coated Carbide tool

20 0

20

40

60

80

100

120

Clearance Wear (Pm) Figure 3.8 Relationship between normal force and clearance wear of the P30 carbide tool and coated carbide tools (in average)

Noise level (dB-C)

113 y = 0.11x + 101.9 r = 0.99

110

y = 0.08x + 102.9 r = 0.83

107 P30 carbide

104

Coated carbide tool

101 0

20

40

60

80

100

120

Clearance Wear µm)

Figure 3.9 Relationship between noise level and clearance wear of the P30 carbide tool and coated carbide tools (in average)

41


Normal Force (N)

50 45 40

y = 3.56x - 345.38 r = 0.99

35 30 25 20 102

104

106

108

110

Noise Level (dB-C) Figure 3.10 Relationship between normal force and noise level of the tools investigated (in average)

3.4 Effect of Work Materials on the Wear and Delamination The results in Figure 3.11 indicate that the amount of clearance wear attained by the uncoated P30 and TiN coated P30 was observed to be greater for the wood-chip cement board than for the hardboard at the same density. The difference in clearance wear between hardboard and wood-chip cement board for the uncoated and TiN coated P30 was more prominent in high-density than in low density. The difference in the high-density was about 60 µm for the uncoated P30 and about 30 µm for the TiN coated P30. This difference is considered to be due to the two following reasons. First, the wood-chip cement board of low or high density contains a large amount of hard particle of cements, which will lead to exert higher abrasion on the cutting tools compared to the hardboard during the cutting process. The hard particles of cement generally consist of limestone, alumina, and silica. Therefore, the harmful effect of the hard particles of cement is considered to be the same as that of silica found in wood [8,9] or that of cured formaldehyde resin in

42


HD Cement board HD Hardboard LD Cement board LD Hardboard

180 150 120 90 60 30 0 0

250 500 750 100 125 150 175 200 0 0 0 0 0

Length of Cut (m)

Clearance Wear (Pm)

Clearance Wear (Pm)

particleboard [10,11], which caused cutting tool to wear out rapidly. Second, the fine particles produced during cutting wood-chip cement board both in low and high density were observed to adhere, embed and drag on the surfaces of the cutting edges (Figure 3.12a–c).Whereas continuous chips (Figure 3.12b–d) produced during cutting hardboard of low and high density escape freely up to the rake face to impose less abrasion. 180


150 120 90 60 30 0 0

250 500 750 100 125 150 175 200 0 0 0 0 0

Length of Cut (m)

Figure 3.11 Clearance wear progress curves of the uncoated P30 (left) and TiN coated P30 (right) in cutting the hardboards and woodchip cement boards. LD = low-density, HD = high density

43


Figure 3.12 The formation of chips and the cutting edges condition of TiN coated P03 in cutting wood-chip cement boards (a-c) and hardboards (b-d). FP (fine particles), CE (cutting edge) It also appears from the results in Figure 3.11 that the clearance wear of the cutting tools increased by increasing in the density of the work materials. This behavior was found to be more remarkable in cutting wood-chip cement board than in cutting the hardboard. The clearance wear is observed to be 55 µm greater in cutting the wood-chip cement board of high density than that of low density for the uncoated P30, and to be 24 µm greater in cutting the wood-chip cement board of high density than that of low density for the TiN coated P30. It could be considered as a reason that more materials are machined at a given cutting volume for the high-density wood-chip cement board and the mechanical strengths are higher in the high-density wood-chip cement board (Table 3.2). This fact might also relate to the phenomenon that the higher the density in the work materials the higher the abrasiveness will be [12], which leads to impose higher abrasion during the cutting. However, only slight differences in clearance wear were observed in cutting hardboard of low and high density using both uncoated P30 and TiN coated P30 (Figure

44


3.11). This could be caused by almost the same compressive strength between them (Table 3.2), though more materials are present in a given cutting volume for the high-density hardboard. Table 3.2 Specifications of the work materials machined Compressive Strength (N/mm2)

Shear Strength (N/mm2)

Hardness (N/mm2)

10.5

3.4

0.6

10.0

8.1

110.8

0.7

33.8

1.20

11.6

23.4

3.5

50.6

1.18

7.4

118.8

4.7

70.9

Thickness (mm)

Density (g/cm3 )

MC (%)

LD Cement board

25

0.79

LD Hardboard

25

0.81

HD Cement board

25

HD Hardboard

25

Work Material

LD = low density, HD = high density, MC (moisture content)

Investigations of the worn edges of the TiN coated P30 under SEM show that relatively the same patterns of tool wear were observed both in cutting hardboards and wood-chip cement boards (Figure 3.13). It is observed from the results in Figure 3.13 that the wear of the carbide substrate and delamination of the TiN film occurred mainly due to mechanical abrasion with small contribution of chemical reaction. The chemical reaction occurred in cutting wood-chip cement board was clearly characterized by the formation of cement compounds (cc) adhere on both the substrate and coating film surfaces (Figure 3.13c–d). Darmawan et al. [13] noted that these adhered cement compounds were investigated to impose high mechanical abrasion on the substrate and coatings, and accelerated the wear of the coated carbide tools. It is considered that the wear of the TiN coated P30 was proceeded by delamination of TiN film due to mechanical abrasion and afterwards the wear of the substrate occurred as the TiN film was disappeared from the substrate.

45


Rake

a

c Sb

Clearance b

50 µm

d

CM

50 µm a = c = Sb cc

cc

TiN-coated for LD Hardboard TiN-coated for LD Cement board = Substrate, CM = Coating Material = cement compound

50 µm

cc

50 µm b = TiN-coated for HD Hardboard d = TiN-coated for HD Cement board AbCM = Abrasion of Coating Material

Figure 3.13 SEM Micrograph of TiN coated P30 at 2 km cutting length for the hardboards and wood-chip cement board The results in Figure 3.14 show that the normal force generated by the uncoated P30 and TiN coated P30 increased in proportion with an increase in cutting length. The uncoated P30 carbide tools, compared to TiN coated tools, showed larger proportion of normal force especially near end of cutting for all materials machined [14]. The higher normal force of the uncoated P30 was caused by higher clearance wear suffered by the uncoated P30. Therefore, the TiN coated P30 is considered to be more advantageous not only in the wear resistance, but also in the normal force than uncoated P30 carbide tool with increasing in cutting length for the hardboards and wood-chip cement boards.

46


Normal Force (N)

50 40 30 20 10 0

60

Normal Force (N)


60

50 40 30 20 10 0

0

250 500 750 10001250150017502000

Length of Cut (m)

HD Cement board HD Hardboard LD cement board LD hardboard

0 250 500 750 10001250150017502000

Length of Cut (m)

Figure 3.14 Normal force behaviors with increasing in cutting length for the uncoated P30 (left) and TiN coated P30 (right). LD = lowdensity, HD = high density It also appears from the results in Figure 3.14 that the normal forces generated by the uncoated and TiN coated P30 were observed to increase by increasing in the density of the work materials. This fact is considered to be due to the effect of higher amount of wear attained by the uncoated and TiN coated P30 in cutting the work materials of high density than that of low density, and due to the higher in strengths of the high density work materials (Table 3.2). Though considerable differences in wear were not observed in cutting the hardboard of low and high density using uncoated and TiN coated P30, however a significant differences in normal force were observed. This fact is probably caused by the difference in shear strength and hardness (Table 3.2) between the low and high-density hardboard. The results in Figure 3.14 show that though the uncoated and TiN coated P30 attained higher amount of wear in cutting low-density wood-chip cement board than that in low-density hardboard, the normal forces generated during cutting low-density wood-chip cement board were much lower. The uncoated and TiN coated P30 generated normal force in cutting low-density woodchip cement board almost half than that in low-density hardboard since beginning of cutting up to cutting length of 2 km [14]. This is considered to be due to the fact that the low-density wood-chip cement board is much more porous in structure with low compressive strength and hardness (Table 3.2) compared to low-density hardboard having more solid structure with much higher in compressive strength and hardness. The results in Figure 3.14

47


also indicate that the normal forces of the uncoated and TiN coated P30 in cutting low density wood-chip cement board appeared to be constant, though its wear and delamination increased with the cutting length. Different phenomenon in normal force was observed when the uncoated and TiN coated P30 cut the work materials of the high density (Figure 3.14). The uncoated and TiN coated P30 generated slightly higher normal force in cutting high-density wood-chip cement board than that of high-density hardboard at the end of cutting (near 2 km of cutting length). However, near beginning of cutting, the uncoated and TiN coated P30 generated slightly lower normal force in cutting high-density wood-chip cement board than that of high-density hardboard. This fact is considered due to the high-density wood-chip cement board being lower in strengths and hardness compared to the high-density hardboard (Table 3.2). Therefore, the behavior of normal force at the end of cutting described above is caused by higher amount of wear attained by the uncoated and TiN coated P30 tools in cutting high density wood-chip cement board than that of high density hardboard.

References 1. Darmawan W, Tanaka C, Usuki H, Ohtani T (2000) Wear characteristics of some coated carbide tools when machining hardboard and wood-chip cement board. J. Wood Industry. Vol. 55 (10): 456–460. 2. Darmawan W, Tanaka C, Usuki H, Ohtani T (2001) Performance of coated carbide tools when grooving wood-based materials: Effect of work materials and coating materials on the wear resistance of coated carbide tools. J. Wood Science. Vol. 47 (2): 94–101. 3. Biernat S (1995) Carbide Coatability, Cutting Tool Engineering. 47(3): 44–45. 4. Smith GT (1989) Advanched Machining–The Handbook of Cutting Technology. Springer-Verlag, London, pp150-200. 5. Hunt JL, Santhanam AT (1990) Coated carbide metal cutting tools: Development and applications. PED-Vol. 43, p139–155. 6. Davis J R (1998) Tool Material. ASM International. USA, pp77–84.

48


7. Darmawan W, Tanaka C (2004) Discrimination of coated carbide tools wear by the features extracted from parallel force and noise level. Ann. For. Sci. 61 (2004) 731–736. 8. Darmawan W, Rahayu IS, Tanaka C (2006) Chemical and mechanical wearing of woodworking cutting tools by tropical woods. Journal of Tropical Forest Science. 18(4): 166–172. 9. Hayashi K, Suzuki T (1983) Effect of cutting speed on tool wear in the peripheral milling of wood (in Japanese). Mokuzai Gakkaishi. 29(1): 36– 42. 10. Kollmann FFT, Kuenzi EW, Stamm AJ (1975) Principles of wood science and technology II (wood-based materials). Springer-Verlag, New York. 11. Kim MG, Stewart HA, Wan H (1999) Effects of Anionic Components of Urea Formaldehyde Resins Used as Particleboard Binders. For Prod J. 49(4): 60–65. 12. Bridges RR (1971) A Quantitative Study of Some Factors Affecting the Abrasiveness of Particleboard. For Prod J. 21(11): 39–41. 13. Darmawan W, Tanaka C, Usuki H, Ohtani T (2001) Performance of coated carbide tools in turning wood-based materials: Effect of cutting speeds and coating materials on the wear characteristics of coated carbide tools in Turning Wood-Chip Cement Board. J Wood Science. 47(5): 342–349 14. Darmawan W, Usuki H, Marchal R, Quesada J. (2008): Clearance Wear and Normal Force of TiN-coated P30 in cutting hardboard and woodchip cement board. Holz als Roh- und Werkstoff. Vol 66(2): 89–97.

49

IV. Performance of Monolayer Coated Tools in High Speed Cutting of Wood-Chip Cement Board 4.1 Effect of Work Materials on the Cutting Temperature Superiority of P30 carbide tool coated with TiN, TiCN, and CrN coating by physical vapor deposition (PVD) method for cutting of hardboards at low cutting speed was discussed in the previous sections. Further, the possibility of the P30 carbide tool coated with the monolayer of TiN, CrN, TiCN, and multilayer of TiN/AlN coatings by PVD method on the rake and clearance faces for higher speed machining (Table 4.1) is discussed in this section. Table 4.1 Specifications of the work materials, tools and cutting conditions Coating material

Thickness of film (m)

Hardness (HV)

Resistance to oxidation temperature (oC)

Absorptive capacity of heat energy (Ws1/2/ m2K)54

P30 carbide Coated carbide

TiN CrN TiCN TiN/AlN

3–4 3–4 3–4 3–4

1450 2000 1800 3000 4000

750 800 450 930

8100 13900 6300

Work material (WM)

Thickness (mm)

Density (g/cm3 )

Moisture content (%)

Compressive strength (N/mm2)

Shear strength (N/mm2)

Hardness (N/m2)

Wood-chip cement board

25

1.20

11.6

23.4

3.5

50.6

Cutting condition

Cutting speed, V, (m/s)

Feed, F, (mm/rev)

Width of Cut, W, (mm)

30, 40, 50, 60

0.05

1

Tool

Tool geometry Wedge angle = 90o, rake angle = -5o, clearance angle = 5o, corner radius = 0.8 mm


Work materials (Table 4.2) are one of the important factors that affect the cutting tool temperature, which leads to the tool wear. The cutting tool temperatures measured by EMF method [1] varied among the work materials tested and evidently also depended on the density of the work materials, as presented in Figure 4.1.It appears that the differences in cutting tool temperature among the work materials were very small at low cutting speed. However, the increase in cutting speed resulted in more remarkable differences in the cutting tool temperature. Table 4.2 Work materials and conditions for the tools temperature measurement Tool couple

Work materials

Cutting conditions

K10 carbide and Cermet (wedge angle of 90o, rake angle of –5o, clearance angle of 5o) Thickness (mm) Density (g/ Moisture cm3 ) content (%) HD CB

5

1.20

11.6

HD HB

5

1.18

8.1

LD CB

5

0.79

10.5

LD HB

5

0.81

7.4

Douglas fir 5 0.32 11.0 Cutting Contact area Cutting Feed, F, Width of speed, V, between the period (mm/rev) Cut, W, (mm) (m/s) tool (mm2) (s) 20, 30, 40, 0.05 5 0.125 3 50, 60

HD = high density, LD = low density, CB (wood-chip cement board), HB (hardboard)

The results in Figure 4.1 show that hardboard and wood-chip cement board of high-density generated higher cutting tool temperature compared to those of low-density at every cutting speed performed. This fact is caused by more material being cut in a given cutting volume for the high-density work materials leading to great friction on the cutting tools. Higher compressive and shear strengths in the high density work materials are also considered to impose great friction on the cutting tools which gives important contribution to the heating of the cutting tools.

52

IV. Performance of Monolayer Coated Tools in High Speed Cutting of Wood-Chip Cement Board

Cutting Temperature (oC)

1200 1000

HD CB

HD HB

LD CB

LD HB

Douglas Beimatsu fir

800 600 400 200 0

20

30

40

50

60

Cutting Speed (m/s) Figure 4.1 Tool temperature as a function of cutting speeds and work materials Comparison of the work materials in Figure 4.1 shows that wood-chip cement board of high-density generated evidently the highest cutting tool temperature. Therefore, this behavior is another reason why the wood-chip cement board of high density was chosen as the work material for determining the wear characteristics of coated carbide tools, which were experimented in this section [1]. The presence of cement particles, which generally contain limestone, alumina and/or silica, in the wood-chip cement board is considered to cause strong friction on the cutting tool compared to wood fibers in hardboard and douglas-fir. This strong friction would be considered to produce much heat during the cutting. Furthermore, fine cement particles produced during turning wood-chip cement board tended to adhere on the cutting tools [1] compared to wood dusts in hardboard and douglas-fir that tended to escape from the cutting tools. These adhered-cement particles would contribute to the strong friction on the tool surfaces, which lead to raise the cutting tool temperature, as they would abrade the cutting tool. Though the cutting tool temperature was the highest for the high-density wood-chip cement board, however it was the lowest for the low-density woodchip cement board. This is due to the fact that the low-density wood-chip cement-board was porous in structure and very low in mechanical strength

53


(Table 4.1) compared to the other work materials. These characteristics would allow less friction between tool and work material, which produced low heat during the cutting.

4.2 Clearance Wear and Cutting Tools Life Figure 4.2 presents the progressions of the clearance wear with cutting time at four different cutting speeds for all tools investigated. Almost the same clearance wear behaviors were observed among the tools investigated. It appears that the differences in the amount of clearance wear among the cutting speeds were not remarkable at the beginning of cutting, however greater differences in the clearance wear were developed and followed by an approximately linear increase in wear with cutting time.Therefore, rates of the clearance wear, which were characterized by the coefficients of regression equation of the linear, were calculated using Figure 4.2, and the results are summarized in Table 4.3. Behaviors of clearance wear of the tools investigated with cutting speeds (V) at 1.3 minute cutting time (T), which were calculated from the linear regression equations in Figure 4.2, are presented in Figure 4.3. The results in Table 4.3 and Figure 4.3 give an indication that higher rates of wear and greater amount of wear, respectively, were provided by almost all tools investigated when higher cutting speeds were performed. The coated carbide tools were better in these values compared to those of the P30 carbide tool. This fact is considered to be due to the coated carbide tools having higher hardness than P30 carbide tool, which resulted in their resistance to mechanical abrasion to be higher than the P30 carbide tool [1]. It is observed that the clearance wear of the carbide substrate are smaller in the coated carbide tools for low cutting speed (30 m/s), than in the coated carbide tools for high cutting speed (60 m/s). It appears that the wear of the tools investigated also occurred by the delamination of coating film in both cutting speeds. However, the extent of delamination between the two cutting speeds is largely different. This fact would suggest that the mechanism of coating film delamination occurred in different manner between the two cutting speeds.

54

400 400

Clearance Wear (m)

Clearance Clearance Wear Wear (m) (m)


300 300 200 200 100 100 00 0.0 0.0

a 0.5 0.5

1.0 1.0

1.5 1.5

2.0 2.0

400 300 200 100 0

2.5 2.5

b 0.0

Clearance Wear ( m)


400 300 200 100

c 0.0

0.5

1.0

1.5

2.0

2.5


Cutting Time (min) 400 300

1.5

2.0

2.5

400 300 200 100 0 0.0

d 0.5

1.0

1.5

Cutting Time (min)

2.0

V 30m/s

V 40m/s

V 50m/s

V 60m/s

2.5

a : P30 carbide

200

b :

100 0 0.0

1.0

Cutting Time (min)

Cutting Cutting Time Time (min) (min)

0

0.5

TiN coated tool

c : CrN coated tool

e 0.5

1.0

1.5

2.0

d : TiCN coated tool 2.5

e : TiN/AlN coated tool

Cutting Time (min)

Figure 4.2 Clearance wear behaviors of the tools investigated with increasing the cutting time at four different cutting speeds

55


Table 4.3 Rate of wear of the tools investigated at four different cutting speeds Cutting speed (m/s)

P30 carbide CrN coated carbide TiN coated carbide TiCN coated carbide TiN/AlN coated carbide

Clearance Wear (µm)

40

50

60

118.0 112.9 76.6 61.6 33.4

167.9 159.8 163.3 164.5 47.4

276.5 241.5 242.7 276.1 52.2

Temperature P-30 carbide TiN/AlN TiN CrN TiCN

500 400 300 200

30 39.0 37.3 35.0 24.2 25.7

1200 1010 800

800

630

600

435

400

100 0

1000

200 30

40

50

60

0

Tool Temperature (oC)

Tools

Cutting Speed (m/s) Figure 4.3 Behaviors of clearance wear with an increase in cutting speed at 1.3 minute cutting time and predicted tool temperature for each cutting speed The SEM/EDS images of the worn edges of the coated carbide tools investigated are presented in Figure 4.4. Though the TiCN coated carbide tool was the lowest both in rate of wear and amount of wear for cutting speed of 30 m/s (Table 4.3 and Figure 4.3), their values increased drastically and matched those of the CrN and TiN coated carbide tools investigated when cutting speed was beyond 50 m/s [1]. Therefore, the TiCN coated

56


carbide tool would be suitable for low cutting speed. This phenomenon is considered to depend on that the TiCN coated carbide tool will be able to retain its high hardness at the low cutting speed, and that it is escaped from oxidation, because the cutting tool temperature was observed to be still under its oxidation temperature (Figure 4.3).

Figure 4.4 SEM/EDS micrographs that show the presence of oxides (red color) on the worn edges of the TiN, TiCN and TiN/AlN coated carbide tools at 1.3 minute cutting time (5 km cutting length) Severe increase of wear of the TiCN coated carbide tool beyond 50 m/s cutting speed is considered to be due to the fact that the cutting tool temperature exceeded its oxidation temperature. This high temperature was observed to cause severe oxidation of TiCN coating, which is observed in the SEM/EDS image in Figure 4.4b. It is observed in that Figure that plentiful oxides (shown in red color) are present on the coating film near delamination line (DL) of the TiCN coated carbide tool. This leads to that the TiCN coating suffered

57


thermal degradation. As the result, severe delamination of TiCN coating was occurred at the high cutting speeds. Further, the hardness of the TiCN coating material was reported to decrease to about 1100 Hv at temperature of 800oC [2]. These suggest that the rate of wear and the amount of wear of the TiCN coated carbide tool were accelerated by the oxidation of its coating film and the decrease in its hardness due to the high temperature. It is known, moreover, that the hardness of the substrate beneath it decreases rapidly at temperature as low as 600 oC [3] and that the TiCN coating has very high absorptive capacity of heat which allows much heat to conduct and/or concentrate to the substrate (Table 4.1). Therefore, the cutting temperatures of about 800 and 1010 oC determined at the 50 and 60 m/s cutting speeds, respectively (Figure 4.3) are considered to be responsible for the damage of the substrate due to severe loss in hardness. The hardness of P grade carbide (10% Co, 84% WC) was reported to be about 600 Hv at 800oC and 500 Hv at 900oC [3]. Same patterns of wear as those in the TiCN coated carbide tool were also observed in the CrN and TiN coated carbide tools (Figure 4.4a), were high both in rate of wear and amount of wear at the 50 and 60 m/s cutting speeds (Table 4.3, Figure 4.3). Therefore, the CrN and TiN coated carbide tools are considered to follow the behaviors of the TiCN coated carbide tool explained above. The CrN and TiN coated carbide tools were also high in these values for cutting speeds of 40 m/s (Table 4.3, Figure 4.3). Though the 630 oC cutting tool temperature for the 40 m/s cutting speed was under their oxidation temperature, however this temperature is considered to decrease the hardness, which weaken its resistance to mechanical abrasion. This would be confirmed by the fact that the hardness of TiN coated carbide tool decreased with an increase in temperature, and its hardness at 600 and 800oC was about 850 and 625 Hv, respectively [4]. On the contrary, these values of TiN/AlN coated carbide tool did not markedly increase with an increase in the cutting speed (Table 4.3 and Figure 4.3). Their values at cutting speed of 60 m/s were almost equal or lower compared to those at 40 m/s of the other tools investigated. This result will suggest that the TiN/AlN coated carbide tool will provide wider ranges of cutting speed in its application [1]. Further, in the viewpoint of the suitability of the substrate,

58


the P30 carbide tool will be probably better choice for the TiN/AlN coating material. This is attributed to that the TiN/AlN film coated on P30 carbide can resist the delamination and/or chipping compared to the other coated carbide tools investigated. The TiN/AlN coating film strongly bound with the carbide substrate beneath it for the 60 m/s cutting speed and protected the substrate from any wear behavior compared to the other coating films investigated. Though the cutting temperature for the 60 m/s cutting speed exceeded its oxidation temperature, however it was observed that almost no oxidation, which will lead to the thermal degradation, occurred on its coating film (Figure 4.4b). This is considered to be due to the TiN/AlN coated tool being composed by the alternating-layers of very thin TiN and AlN film (20 nm each), in which the AlN coating provided a barrier which resisted chemical interaction [4]. Moreover, the TiN/AlN coated carbide tool is very high in hardness and low in absorptive capacity of heat energy (Table 4.1). Because of this superiority, it retains heat in the surfaces and retard the conduction of the heat to the substrate. This behavior will lead to the retention of high hardness of the substrate and will protect the substrate from the wear. Therefore, the coating materials that will be synthesized on the substrate should be high in hardness, high in the oxidation temperature and low in absorptive capacity of heat energy (low heat conductivity). Choosing a given tool life criterion of 200 µm clearance wear, a corresponding tool life time for each cutting speed can be found in Figure 4.2 and is plotted in log scales as shown in Figure 4.5. On the log cutting time (T) – log cutting speed (V) plots, the tool life exponent n and constant C in the Taylor tool life equation was calculated. The results are shown in Table 4.4.

59

Cutting Speed ,V, ( m/s)


60 50 40 Uncoated tungsten TiAlN TiN CrN TiCN

30

0.5

1.0

2.0

3.0

5.0

10.0

Cutting Time ,T, (min) Figure 4.5 Tool life and cutting speed relationship for the tools investigated Table 4.4 The exponent n and constant C of the Taylor tool life equation (VTn=C) [1] Tools P30 carbide CrN coated carbide TiN coated carbide TiCN coated carbide TiN/AlN coated carbide

Exponent n 0.33 0.34 0.36 0.29 0.73

Constant C 47.3 49.6 54.6 54.7 132.0

Experimental results presented in Table 4.4 show that the values of n vary slightly among the tools investigated. It is well known that the Taylor tool life model is used to explain the difference of wear mechanism. It is known that tool life exponent n having a value of unity presents relatively same wear mechanism, in which value of n equal one determines the wear of cutting tool due to the mechanical abrasion. Therefore, the values of tool life exponent for cutting wood-chip cement board shown in Table 4.4 will give an indication that the tools investigated underwent slightly different wear mechanisms. The values of n of the TiN/AlN coated tool in Table 4.4, which is approximately close to one, will suggest that its wear was determined mainly by mechanical abrasion and very small contribution of chemical reaction (oxidation), as discussed above. It appears that the life of the P30 carbide, TiCN, TiN,

60


and CrN coated carbide tools was strongly affected by the cutting speeds or indirectly by the cutting tool temperature (Figure 4.5). This phenomenon gives an indication that though they probably suffered the same mechanical abrasion as the TiN/AlN coated carbide tool, however they suffered more severe chemical reaction (oxidation) [1]. The results in Table 4.4 indicate that TiCN coated carbide tool suffered the largest chemical reaction followed by the P30 carbide, CrN and TiN coated carbide tools. Further, comparison of the value of C in Table 4.4 indicates that the TiN/AlN coated carbide tool will be the longest in tool life and the P30 carbide tool will be the shortest. Specifically the cutting speed for 1 minute tool life for the TiN/AlN coated carbide tool is predicted to be 132.0 m/s followed by TiCN (54.7 m/s), TiN (54.6 m/s), CrN (49.6 m/s) and P30 (47.3 m/s).

4.3 The Effect of Cutting Speed on Forces and Noise Some automatic monitoring systems have been proposed to achieve a higher productivity in wood processing. Application of the automatic monitoring system in band-sawing, circularsawing, routing, milling and peeling would help the wood worker in increasing productivity, diagnosing the machine conditions (bearings, chains), predicting the machining imperfectness (washboard, abnormal roughness), and controlling the cutting. tool condition (tool edge wear, tool edge damage). Several parameters (cutting energy [5], cutting force [6], cutting sound [7], acoustic emission [8], saw vibration [9], cutting temperature [10]) were investigated to be useful for providing information to the automatic systems. Cutting forces and noise level have shown great promise in the monitoring of the extent of tool wear. An excellent correlation was found to exist between the cutting forces and tool wear [11, 12, 13], and the feasibility of techniques of pattern recognition using the cutting sound for the discrimination of the various stages of edge wear was clarified [14]. Another study revealed that behavior of acoustic emission (AE) signal was found to be a superior parameter for estimation of wear of the cutting tool [15]. In those studies, non-coated tools with their various geometries were experimented.

61


In another study, coated carbide tools were experimented for cutting wood based materials, and their wear characteristics were reported [16,17]. In an effort to provide more information, the parallel force and noise level of the tools, and their changes due to the wear were reported in this paper. The regression equations were applied to discriminate the various stages of tool wears by the features extracted from the parallel force and noise level. The purpose is to determine the feasibility of using parallel force and noise level to monitor the extent of wear of the coated carbide tools for various cutting speed. Figure 4.6 shows the variation of cutting forces with increasing the cutting speed at 1.3 minutes cutting time. It appears that both the feed (Figure 4.6a) and parallel forces (Figure 4.6b) increased linearly with an increase in cutting speed for almost all tools investigated. The results in Figure 4.6 give an indication that the feed force showed similar behaviors as the parallel force. However, evidently the parallel force appears to be higher than the feed force at every cutting speed performed. The increase in cutting force with an increase in cutting speed could be caused by more work material being machined within a given cutting time and greater amount of wear being attained by the cutting tool edge for the high cutting speed (Figure 4.3). Experimental results in Figure 4.6 show that the effect of coating materials on the feed and parallel forces is very small. The TiN/AlN coated carbide tool generated the smallest both feed and parallel forces for all cutting speeds among the tools investigated [18]. This is considered to be due to the fact that the TiN/AlN coated carbide tool suffered much lower amount of wear at every cutting speed performed compared to the other tools investigated, as given in Figure 4.3. Though the other tools investigated were relatively same in the feed force, they varied slightly in the parallel force as the cutting speed was increased.

62


P30 carbide TiN TiCN CrN TiN/AlN

32 24

40

Parallel Force (N)

Feed Force (N)

40

16 8 0

a 25

35

45

55

65

b

32 24 16

P30 carbide TiN TiCN CrN TiN/AlN

8 0

25

35

45

55

65

Cutting Speed (m/s)

Cutting Speed (m/s)

Figure 4.6 Behaviors of feed force (a) and parallel force (b) of the tools investigated with cutting speed at 1.3 minute cutting time P 30 CrN TiN/AlN TiCN TiN

Noise level (dB-C)

94 93 92 91 90

20

30

40

50

60

70

Cutting speed (m/s) Figure 4.7 Noise level behaviors with cutting speed at 1.3 minute cutting time for all tools investigated Figure 4.7 shows the variation in noise level with an increase in cutting speed at 1.3 minutes cutting time. It appears that the noise levels also increased linearly with an increase in cutting speed for almost the all tools investigated. The high noise levels generated during high speed cutting are probably also caused by more work material being machined in a given cutting time and large impact force being imposed on the tools for the high speed cutting [18,19]. It is observed from the results in Figure 4.7 that though the amount

63


of wear of the TiN/AlN coated carbide was much lower than the other tools investigated, however its noise level is almost the same as that of the other tools. The results in Figure 4.8a indicate that the noise level and feed force are close in relationship. The feed force increased linearly with increasing the noise level. It appears that the cutting speed influenced to their relationships. When the noise level increased in 3.0 dB, the feed force of the tools increased about 3.2 N for 30 m/s cutting speed, and about 8.9 N for 60 m/s cutting speed. The results in Figure 4.8b also indicate that the noise level and the parallel force are close in relationship. The noise level increased linearly with increasing the parallel force. It could be considered from the regression equation in Figure 4.8b that the tools (in average) would generate about 1 dB noise level when the parallel force of the tools changed in 2 N for the 30, 40, and 50 m/s cutting speeds, and would generate about 1 dB noise level when the parallel force changed in 4 N for the 60 m/s cutting speed. This fact gives an indication that the feed and parallel forces became more sensitive compared to the noise level in determining the clearance wear behavior of the tools investigated when the cutting speed was increased [18].

64


35

30m/s 50m/s

Feed Force (N)

30

40m/s 60m/s

y = 2.98x - 228.92 r = 0.88

y = 2.96x - 246.18 r = 0.98

25 20

y = 2.66x - 220.43 r= 0.96

15 10 5

a

y = 1.06x - 79.23 r = 0.84

86

88

90

92

94

Noise Level (dB-C)

95

Noise level (dB-C)

30, 40, 50 m/s (average) 60m/s

93 91

y = 0.49x + 79.65 r = 0.99 y = 0.25x + 83.94 r = 0.94

89 87 85

b 10

15

20

25

30

35

Parallel force (N) Figure 4.8 Relationship between feed force (a), parallel force (b) and noise level of the tools investigated (in average) for four different cutting speeds

65


References 1. Darmawan W, Tanaka C, Usuki H, Ohtani T (2001) Performance of coated carbide tools in turning wood-based materials: Effect of coating materials and cutting speeds on the wear characteristics of coated carbide tools in turning wood-chip cemenboard. J. Wood Sci. 47 342–349 2. Wysiecki M (1997) Modern Cutting Tool Materials.Wydawnictwa Naukowo-Techniczne. (in Poland Language), Poland. p50–55. 3. Millman YV, Luyckx S, Northrop IT (1999) Influence of temperature, grain size and cobalt content on the hardness of WC-Co alloys. International Journal of Refractory Metals & Hard Materials. 17(1999)39– 44. 4. Hunt JL, Santhanam AT (1990) Coated carbide metal cutting tools: Development and applications. PED-Vol. 43, p139–155. 5. Piotr I, Tanaka C (2003) Energy Balance during orthogonal machining of medium density fiberboard, in: Proceedings of the 16th International Wood Machining Seminar, pp. 459–467. 6. Eyma F, Meausoone PJ, Martin P (2004) Study of the properties of thirteen tropical wood species to improve the prediction of cutting forces in modes B. Ann. For. Sci. Vol. 61: 55–64. 7. Ying-jie Q, Zhao-hao Z, Xiao-jie Q, Shou-qian C, Li Z (2003) Noise measuring and analysis study of precision panel saw, in: Proceedings of the 16th International Wood Machining Seminar, pp. 696–701. 8. Lemaster R, Tee L (1985) Monitoring tool wear during wood machining with acoustic emission, Wear, 101: 273–282. 9. Ulsoy AG, Mote CD Jr (1982) Vibration of wide band saw blades. J. Engin. Ind. Trans. ASME. 104: 71–78. 10. Sokojowski W, Gogolewski P (1999) Temperature of machined surface as value for tool condition monitoring during wood products milling, in: Proceedings of the 14th International Wood Machining Seminar, pp. 775–780. 11. Axelsson BOM, Grundberg SA, Gronlund JA (1993) Tool wear when planning and milling, Measurement methodology and influencing factors, in: Proceedings of the 11th International Wood Machining Seminar, pp.159–176. 66


12. Tanaka C, Takahashi A, Shiota Y (1986) Cutting performance of cermet, ceramic, CBN and artificial diamond I: Wear from continuous cutting of wood-based materials. Mokuzai Gakkaishi. 32: 96–102. 13. Tanaka C, Takahashi A, Date H, Nakao T (1986) Cutting performance of cermet, ceramic, CBN and artificial diamond. III. Cutting performance of ceramic tools. Mokuzai Gakkaishi. 34: 298–304. 14. Fujii Y, Fuketa T, Arashi Y, Okumura S, Noguchi M (1993) Pattern recognition of cutting sound from woodworking tools and its application to the in-process monitoring of wear, in: Proceedings of the 11th International Wood Machining Seminar, pp. 147–156. 15. Tanaka C, Nakao T, Nishino Y, Hamaguchi T, Takahashi A (1992) Detection of wear degree of cutting tool by acoustic emission signal. Mokuzai Gakkaishi. 38: 841–846. 16. Darmawan W, Tanaka C, Usuki H, Ohtani T (2000) Wear characteristics of some coated carbide tools when machining hardboard and wood-chip cement board. J. Wood Ind. 55: 456–460. 17. Darmawan W, Tanaka C, Usuki H, Ohtani T (2001) Performance of coated carbide tools in grooving wood-based materials: Effect of coating materials and work material on the wear resistance of coated carbide tools. J. Wood Sci. 47: 94–101. 18. Darmawan W, Tanaka C (2004) Discrimination of coated carbide tools wear by the features extracted from parallel force and noise level. Annals of Forest Science. Vol. 61 (7): 731–736. 19. Costes JP, Larricq P (2002) Towards high cutting speed in wood milling, Ann. For. Sci. 59: 857–865.

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V. Wear Characteristics of Multilayer Coated Cutting Tools in Milling Particleboard The previous chapters indicate that the monolayer coatings did not provide significant improvement in the cutting tool life for high speed cutting of wood based materials. Ongoing research efforts have been proposed to achieve better performance of the coated carbide tools in cutting wood-based materials. Multilayer coatings would be a promising technique to improve the performance of the monolayer coatings [1,2,3,4]. TiAlN coating, which is high in hardness, good in oxidation resistance and better in wear resistance among the other monolayer coatings, was multilayered with the newest generation coatings of titatnium boron nitride (TiBON), titanium silicon nitride (TiSiN), and chromium aluminum nitride (CrAlN), which have been noted to keep excellent properties (high hardness, low friction coefficient, high oxidation and corrosion resistances) [5,6]. In this chapter, TiAlN coating that multilayered with TiSiN, TiBN, and CrAlN coatings onto the surface of K10 tungsten carbide using the PVD method is discussed. The newly multilayer-coated tools (TiAlN/TiBN, TiAlN/TiSiN, TiAlN/CrAlN) have been experimentally investigated for the possibility of their application in machining particleboard.

5.1 Delamination Wear Delamination wear behavior on clearance faces of the coated tools are provided in Figure 5.1. The results in Figure 5.1 indicate that the amount of delamination wear increased with increasing in cutting length. The multilayercoated tools provided better performance especially in reducing the progression of delamination wear than the monolayer TiAlN in cutting the particleboard. Though the monolayer TiAlN and multilayer-coated tools showed almost the same delamination progress near the beginning of cutting, however the


delamination wear of the monolayer TiAlN increased markedly and exceeded the delamination wear of the multilayer-coated tools, which showed only gradual progression of delamination during cutting the particleboard. The monolayer TiAlN cutting tool suffered delamination wear of about 122 µm for cutting the outer part of the particleboard and about 60 µm for the inner part of the particleboard at the 1 km cutting length. On the other hand, the delamination wear of the multilayer-coated tools was less than 100 µm for the outer part of the particleboard and less than 50 µm for the inner part of the particleboard at the 1 km cutting length. Further, the highest rate of delamination wear occured with the monolayer TiAlN followed by multilayer TiAlN/TiBN and TiAlN/TiSiN (Table 5.1). The lower in hardness, lower in oxidation resistance, and higher in friction coefficient of the monolayer TiAlN compared to the multilayer coated tools (Table 5.2) would be the reason for this phenomenon. a

b

Figure 5.1 Delamination wear behavior of the newly coated tools with cutting length in milling inner part of the particleboard (a) and milling outer part of the particleboard (b) Table 5.1 Rate of delamination wear (µm/km) of the coated tools Bits Outer part Inner part

70

TiAlN 124.7 68.9

Coated cutting tool TiAlN/TiBN TiAlN/TiSiN 108.9 95.3 56.4 51.2

TiAlN/CrAlN -

V. Wear Characteristics of Multilayer Coated Cutting Tools in Milling Particleboard

Table 5.2 Specifications of coated carbide tools tested

28 44 35

Oxidation temperature [8] start at 600oC start at 700oC start at 700oC

Friction coefficient [9] 0.91 0.56 0.61

38

start at 800oC

0.45

Coating material

Film thickness (µm)

Hardness (GPa) [7]

TiAlN TiAlN/TiBN TiAlN/TiSiN

3 3 3

TiAlN/CrAlN

3

It also appears from the results in Figure 5.1 that the outer part of the board (Figure 5.1b) caused higher amount of delamination wear compared to the inner part of the board (Figure 5.1a). The outer part of the particleboard caused delamination of TiAlN, TiAlN/TiBN, and TiAlN/TiSiN coatings 1.5X faster than the inner part of the board (Table 5.1). This behavior was found to occur both in the monolayer TiAlN and the multilayercoated tools (TiAlN/TiBN and TiAlN/TiSiN). The outer part of particleboard usually contains higher dosage of thermosetting resin (resolcinol or urea-formaldehyde in most cases) than the inner part. This is one of the main reasons for the increased tool wear when milling the outer part of particleboard.

Outer part

Surface of board

6 mm

Inner part

Figure 5.2 Photograph showing the structure of the particleboard along the thickness Furthermore, it is depicted in Figure 5.2 that the structure of the inner part of the board was more porous than that of the outer part. It could be considered as another reason that more materials are machined at a given cutting volume for outer part of the board. The ash and silicate contents on the outer part

71


were larger than that on the inner part of the board (Table 5.3). The higher silicate content in the outer part imposed higher abrasion during the cutting. However, no delamination wear was observed for the multilayer TiAlN/ CrAlN in cutting both the outer and inner parts of the particleboard. Table 5.3 Specifications of the particleboard machined Characteristics Thickness Moisture Content Density: Average Outer part Inner part Ash and silicate contents Average Outer part Inner part

Value 12 mm 8% 0.61 (g/cm3 ) 0.73 (g/cm3 ) 0.51 (g/cm3 )

12.15%* 16.56%* 8.61%*

1.86% 2.65% 1.24%

Note : Ash and Silicate content measured according to TAPPI T211 om-85 [10] * Ash content

The results in Figure 5.1 and Figure 5.3 indicate that the multilayer TiAlN/ CrAlN has the highest resistance in delamination wear compared to the multilayer TiAlN/TiSiN and TiAlN/TiBN in cutting the particleboard. The TiAlN/CrAlN did not suffer any delamination of the coating film up to cutting length of 1000 m [11]. The cutting edge of the TiAlN/CrAlN suffered slight chipping of the coating film at the cutting length of 800 m, and still retained the slight chipping up to the final cutting length without any delamination of coating film (Figure 5.3).

72


Figure 5.3 Wear patterns of the coated cutting tools before cutting (a) and after 1 km cutting length (b) The high delamination wear resistance of the TiAlN/CrAlN is considered to be due to the two following reasons. First, the friction coefficient of the multilayered TiAlN/CrAlN coatings was lower than that of TiAlN, TiAlN/ TiBN, and TiAlN/TiSiN, which lead to less abrasion against the hard abrasive materials contained in the particleboard. Other studies reported that a coating material with the highest coefficient of friction shows the highest surface roughness, and further affect life time of cutting tools [12-13]. The 73


hard abrasive materials in the particleboard consisted of mainly cured resin and silicate (Table 5.3), which generate harmful effect on the edge of the coated cutting tool. Second, oxidation temperature data in Table 5.2 and EDS results in Figure 5.4a–b, which were produced by mapping oxygen element from SEM micrograph of edges of the coated tools before and after cutting, suggest a better oxidation resistance retained by the multilayered TiAlN/ CrAlN coatings compared to the TiAlN, TiAlN/TiBN, and TiAlN/TiSiN. Higher peaks of oxygen profiles depicted by TiAlN, TiAlN/TiBN, TiAlN/ TiSiN on the EDS analysis in Figure 5.4b indicates the occurrence of more severe oxidation on these coatings. This phenomenon could be confirmed by a previous result in which delamination of monolayer coatings (TiN, TiAlN, TiCN, CrN) in high speed cutting of wood chip cement board is caused by higher contribution of oxidation [14]. The oxidation was reported to be accelerated by the increase in cutting temperature up to 800o C due to increase in cutting speed above 30 m/s. A possible high temperature generated during high speed cutting of particleboard in this experiment would oxidize the TiAlN, TiAlN/TiBN, TiAlN/TiSiN coated carbide tools, which lead to the severe delamination of coating films due to thermal degradation. TiAlN

TiAlN

TiAlN/TiBN TiAlN/TiBN TiAlN/TiSiN TiAlN/TiSiN TiAlN/CrAlN

TiAlN/CrAlN

(a)

150 µm Oxygen profiles before cutting

(b)

150 µm Oxygen profiles after 1 km cutting l h

Figure 5.4 SEM/EDS analysis showing oxidation indication on the cutting edge of the coated tools before cutting (a) and after 1 km cutting length (b)

5.2 Wear Mechanism Investigations of worn edges of the coated tools under an optical video microscope show that similar delamination mechanisms of coating films were observed in cutting the particleboard. The results in Figure 5.5 depict the 74


mechanism of delamination of TiAlN/TiSiN coatings which is selected for discussion in this article. Delamination of the TiAlN/TiSiN coatings was proceeded by slight chipping of coating film at the cutting edge. The extent of chipping was investigated to increase at the 200 m of cutting length. As the cutting continued up to cutting length of 300 m, the cutting edge underwent more prominent chipping of coating film. Chipping of coating film occurred on the whole cutting edge as the cutting length reached 400 m. Further, the TiAlN/TiSiN coating films were gradually delaminated in proportion along the cutting edge as the cutting action was continued above 400 m. It is considered that the wear of the K10 substrate occurred as the TiAlN/TiSiN films were removed from the substrate. initial edge 0m chipping 100primature m 200 m 300 msevere chipping 400 m delamination 600 m 800 m

K10 1000 m

Figure 5.5 Wear mechanism of the TiAlN/TiSiN multilayered coating in cutting the particleboard SEM micrographs of the worn edges of the TiAlN, TiAlN/TiSiN, TiAlN/ TiBiN, and TiAlN/CrAlN coated tools were presented in cutting the particleboard for 1 km cutting length (Figure 5.6). The SEM micrographs reveal that relatively the same patterns of edge wear were generated by TiAlN, TiAlN/TiSiN and TiAlN/TiBiN [11]. Though severe delaminations were

75


generated by the TiAlN, TiAlN/TiSiN, and TiAlN/TiBiN coated carbide tools, delamination did not occur in the TiAlN/CrAlN coated carbide tool. This fact suggests that the substrate of the TiAlN, TiAlN/TiSiN, and TiAlN/ TiBiN coated tools would be exposed to any possible mechanical abrasion, which caused retraction of tungsten carbide grains from the substrate during the cutting (Figure 5.6a–c). Retraction of carbide grains was reported to cause a corrugated cutting edge, which tends to produce rough quality of board surfaces during the cutting [15,16,17]. Otherwise, TiAlN/CrAlN coatings strongly covered and protected the K10 carbide substrate from wearing, and maintained sharp cutting edge (Figure 5.6d). Cutting edge

1000x

(a) TiAlN

K10 substrate

1000x

(b) TiAlN/TiBN

Chipping 1000x

(c) TiAlN/TiSiN

1000x

(d) TiAlN/CrAlN

Figure 5.6 SEM micrograph of worn edges of the newly coated cutting tools in cutting particleboard at the final cutting length

References 1. Pinheiro D, Vieira MT, Djouadi MA (2009) Advanteges of depositing multilayer coating for cutting wood-based products. Surface & Coating Technology. 203: 3197–3205. 2. Warcholinski B, Gilewicz A (2002) Multilayer coatings on tools for woodworking. Wear. 271: 2812–2820, 2011. Son MJ, Kang SS, Lee EA, Kim KH. Properties of TiBN coating on the tool steels by PECVD and its applications. Journal of Materials Processing Technology. 130–131: 266–271. 76


3. Yang SM, Chang YY, Lin DY, Wang DY, Wu W (2008) Mechanical and tribological properties of multilayerd TiSiN/CrN coating synthesized by a chathodic arc deposition process. Surface and Coating Technology. 202: 2176–2181. 4. Gilewicz A, Warcholinski B, Myslinski P, Szymanski W (2010) Antiwear multilayer coatings based on chromium nitride for wood machining. Wear. 270: 32–38. 5. Ding XZ, Tan ALK, Zeng XT, Wang C, Yue T, and Sun CQ (2006) Corrosion resistance of CrAlN and TiAlN coatings deposited by lateral rotating cathode arc. Thin Solid Films, Volume 516: 987–992. 6. Chang CL, Chen WC, Tsai PC, Ho WY, and Wang DY (2007) Characteristics and performance of TiSiN/TiAlN multilayers coating synthesized by cathodic arc plasma evaporation. Surface and Coatings Technology. 202 (7): 987–992. 7. ASTM (2009) Annual Book of ASTM standard : Standard Test Method for Measurement of Coating Thickness by X-Ray Spectrometry (B568 – 98). Volume 02.05. American Society for Testing and Materials. Philadelphia 8. ASTM (2009) Annual Book of ASTM standard : Standard Practice for Instrumented Indentation Testing (E2546–07). Volume 03.01. American Society for Testing and Materials, Philadelphia. 9. ASTM (2010) Annual Book of ASTM standard : Standard Test Method for Wear Testing with a Pin-on-Disk Apparatus (G99–05). Volume 03.02. American Society for Testing and Materials, Philadelphia 10. TAPPI (1991) Tappi Test Methods: Ash in Wood and Pulp (T211 om85). Volume 1. Tappi Press, Atlanta. 11. Darmawan W, Usuki H, Gottlöber C, Marchal R (2011) Wear Characteristics of Multilayer Coated Cutting Tools in Milling Particleboard. Forest Products Journal. Vol 60 (7/8): 615–621. 12. Venci A, Katavic B, Markovic D, Ristic M, Gligorijevic B (2015) The tribological performance of hardfaced/thermal sprayed coatings for increasing the wear resistance of ventilation Mill working parts. Tribology in Industry. 37(3): 320–329.

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13. Wu W, Chen W, Yang S, Lin Y, Zhang S, Cho TY, Lee GH, Kwon S C (2015) Design of AlCrSiN multilayer and nanocomposite coating for HSS cutting tools. Applied Surface Science. 351: 803–810. 14. Darmawan W, Tanaka C, Usuki H, and Ohtani T (2001) Performance of coated carbide tools in turning wood-based materials : Effect of cutting speeds and coating materials on the wear characteristics of coated carbide tools in Turning Wood-Chip Cement Board. J Wood Science. 47(5): 342–349. 15. Darmawan W, Quesada J, Rossi F, Marchal R, Machi F, and Usuki H (2009) Performance of laser treated AISI-M2 cutting tool for peeling beech. Eur. J. Wood Prod. 67: 247–255. 16. Sheikh-Ahmad J Y, and Bailey JA (1999) High-Temperature wear of cemented tungsten carbide tools while machining particleboard and fiberboard. J Wood Science. 45(6): 445–455. 17. Stewart HA (1992) High-Temperature halogenation of tungsten carbide cobalt tool material when machining medium density fiberboard. For. Prod. J. 42(10): 27–31.

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VI. Characteristics of Laser-Melted T1 High Speed Steel and Its Wear Resistance Resistance of cutting tool materials to wearing is a primary concern in the applicability of tool materials to a cutting operation. The cutting tool material must have adequate resistance to both mechanical and chemical wear. Therefore, the chemical inertness of the tool material, as well as its hardness, must be considered. Different cutting tool materials vary greatly in wear resistance and toughness. Steel cutting tools are the toughest materials, however their relatively low wear resistances limit their application for lower speed machining operation [1]. When using the steel tools for high speed and feed machining of wood and/or wood-based materials, they are required to be high not only in the toughness but also in the wear resistance. The wear resistance of such steels can be increased by employing surface coatings. The hard coatings deposited by chemical vapor deposition (CVD) and physical vapor deposition (PVD) processes have been applied for increasing the performance of the high speed tool steels and experimented for machining woods and wood-based materials. However preliminary delamination of the hard coatings and chipping of cutting tool edges were observed [2,3,4]. Developments made in laser technology over the past 20 years have enabled laser processing an established activity in industry [5]. It was noted in a study that laser application directed to improving the surface properties of materials involve surface melting and surface alloying [6,7]. The application of laser melting on the surface layer of tool materials permits to obtain structures and properties of practical interest. The laser melting permits local thermal treatment for the production of bi-structural tools on the surfaces during a very short time. This laser technology causes the microstructures and properties of the melted surface different from those of its conventional heat treatment.


Laser surface melting is a very promising technique to improve the hardness and microstructures of tool steels with retaining the good toughness of the tool-bulk. It was reported that microstructure of melted AISI-M42 completely changed and its microhardness increased up to 1100 HV compared to its conventional microhardness of 720HV [8]. It was also noted in another study that a fine microstructure was produced after a sintered AISI-42 melted using a laser beam [9]. The hardness of the melted sintered AISI-42 after tempering was reported to increase from 626HV to 820HV. Wear resistance tests of melted steel tools have been already performed and the results were published in a few scientific journals. The increase in the wear resistance of a melted AISI-M2 was reported due to improvement in hardness, in corrosive, erosive and fatigue resistance, and in coefficient of friction [10]. The life of laser melted AISI-M35 bit for cutting a steel material was to be 20% to 125% higher than if the bit was conventionally hardened [11]. The wear resistance of laser-treated Ti—6Al—4V in simulated body fluid was also reported to be enhanced compared to that of the untreated one [12]. In a recent study, an intrinsic surface modification of a steel tool (H13) undertaken by melting using a tungsten arc heat source under nitrogen atmosphere was reported [13]. The best wear resistance was achieved when surfaces were melted under a pure nitrogen or nitrogen–hydrogen shield. In another study, wear resistance of melted steel and plasma coated steel was investigated. The experimental result showed that the wear and corrosion resistances of laser remelted nickel and chromium samples were better than that of plasma coatings [14]. It could be summarized that the changes in microstructure and the increases in hardness due to laser treated as described above, which lead to the increase in the wear and corrosion resistances of the laser-melted steels, were attributed to formation of the very fine microstructure, small proportion of retained austenite, and a higher precipitation of fine carbides within martensite. AISI-T1 (Table 6.1) is one of the high speed tool steel grades, which is traded in the market in form of drills, cutters and end mills. It is high in rupture strength and moderate in toughness, but is relatively low in wear resistance for cutting wood. In this chapter, characteristics of AISI-T1 melted under various laser treatments (Table 6.2) and its performance for cutting wood (Table 6.3) are discussed.

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VI. Characteristics of Laser-Melted T1 High Speed Steel and Its Wear Resistance

Table 6.1 Specifications of AISI-T1 high speed steel for investigation Steel tools for investigation

Dimension of bar (mm)

Conventional Hardened AISI-T1

500 x 40 x 10

Chemical Composition (%wt) 0.78%C, 0.40%Si, 0.40%Mn, 0.03%P, 0.03%S, 4.50%Cr, 1.20%V, 18.00%W

Heat Treatments Austenitizing

Tempering

One hour at temperature of 1230oC

Three time one hour each at temperature of 550oC

Table 6.2 Laser treatment conditions Laser power (watt) 3000, 4000, 5000

Scanning speed (mm/minute) 200, 400, 600

Laser spot diameter (mm) 4

Tempering Three times one hour each at: 525o, 550o, 575o, 600o

Table 6.3 Woods specifications and peeling conditions Wood specification Wood species Density (g/cm3) Moisture content (%) Thickness of disk (mm) Diameter of disk (mm)

Beech 0.71 52 10 400

Peeling condition Cutting speed (m/s) Feed (mm/rev) Thickness of cut (mm) Cutting geometry

1 0.1 0.5 Sharpness angle 25o Clearance angle 1o

6.1 Effect of Laser Treatments on the Depth of Melted Zone The depth of the melted zone on the cross sections of melted surface of T1 sample under different power and scanning speed are presented in Figure 6.1 and Figure 6.2, respectively. The results in these figures show that the depth of melted zone depends on the laser power and the scanning speed. The depth of the melted zone was increased, which means increasing in the melted volume on the substrate, with increasing the laser power (Figure 6.1). Higher power density in watt per unit area generated by the laser beam levelled at the higher power is the reason for producing the larger melt volume. It was noted [15] that power density achieved by laser beam levelled at 0.5–15 kW with 0.1–0.3 mm focusing diameter is in the order of 105–108 Wcm-2 . The melted zone was observed to be shallower when the scanning speed was increased (Figure

81


6.2). This phenomenon is considered due to the decrease in the amount of heat per unit area imposed in the surface of sample, when the scanning speed is increased. MZ

3 kW

Crack

HAZ

S

4 kW

Crack

5 kW

Pore

Figure 6.1 The depth of melted surface under different laser power. Note: speed of 200 mm/min, melting zone (MZ), heat affected zone (HAZ), substrate (S) [16] It also appears from the results in Figure 6.1 and 6.2 that different zones were observed in the cross section of the melted T1 sample. The zones are melted zone (MZ), heat affected zone (HAZ), and substrate (S). The results of investigation, which were conducted on the cross section of the melted T1 under higher magnification, revealed the presence of transition zone. The first transition named TZa is a zone between MZ and HAZ. This zone was a partially melted zone and was considered to correspond with liquidus and solidus interval of the solidification temperature. The second transition named TZb is a zone between HAZ and substrate (S). Fine cracks in the MZ was revealed under laser power of 4000 watt and 5000 watt with scanning speed of 200 mm/min, and pore in the MZ was also revealed under laser power of 5000 watt with scanning speed of 200 mm/min (Figure 6.1) [16]. 200mm/min

400mm/min

600mm/min

TZa TZb

Figure 6.2 The depth of melted surface under different scanning speed. Note: power of 3000 watt, TZa (transition between MZ and HAZ), TZb (transition between HAZ and Substrate)

82


6.2 Microstructures of the Melted Zone (MZ) SEM microstructures of the unmelted T1 (conventionally hardened T1) and the melted T1 are shown in Figure 6.3. Metallurgical structure obtained by conventional heat treatment is a ferritic polycrystalline with coarse primary carbides (Figure 6.3a). Carbides resulting from tempering treatment were not revealed because their sizes are lower than limit resolution of optical and SEM microscopes used in this work. The result in Figure 6.3b shows the microstructure of the MZ, in which whole primary carbides were completely dissolved during laser melting. The microstructure of the MZ was observed to reveal fine iron dendritic structure. It was observed under nital etching that these dendritic structures revealed martensites (dark), as shown in Figure 6.3d. The martensite phase was induced from austenite by a very high cooling speed during the auto quenching of the laser treatment. Fine carbide networks (whitish in colour) formed due to dissolution of the primary carbide grains during laser melting were also observed in the MZ microstructure (Figure 6.3b). Investigation of the MZ microstructure under SEM-EDS reveals that the carbide networks (interdendritic zone) presented lamellar eutectic structure composed of Fe3W3C carbides (Figure 6.3f). This structure resulted from chemical segregation induced by the high cooling speed [16]. The elemental compositions of the dendrite and the interdendritic zone are showed in Table 6.4. It appears that the carbides composed both the interdendritic and dendrite in different proportion.

83


a

b

MZ

c

d

Partially melted carbide TZa

MZ

e

TZa

Fish bond shape carbide

f

MZ

Interdendritic zone

Figure 6.3 SEM microstructures of the unmelted T1 (a), MZ (b, d, f), and TZa (c, e). Note: power of 3000 watt, scanning speed of 200 mm/min, MZ and Tza without tempering [16]

84


Table 6.4 Elemental compositions of the dendrite and the interdendritic zone Element W Cr V Mn

Interdendritic zone 200 mm/min 600 mm/min 27.2 27.3 5.8 5.7 1.5 1.2 0 0.02

Dendrite 200 mm/min 600mm/min 14.5 14.5 3.37 3.6 0.98 0.66 0.2 0.04

Larger iron dendritic crystals, which are also surrounded by carbides, were observed in the TZa (Figure 6.3c). In this transition zone, partially melted primary carbides were visible. They are visible as bright precipitates of typical fish bone morphology and their sizes were observed around 10–20 µm (Figure 6.3e). Under the SEM-EDS analysis the type of carbide in Figure 6.3e was identified as M6C. The M6C carbide was predominantly reach in tungsten (W), vanadium (V) and chromium (Cr). The SEM-EDS microstructures of the MZ without tempering and with tempering were presented in Figure 6.4. The SEM results of the untempered MZ showed no significant difference in microstructure appearance compared to the tempered MZ. However, the EDS analysis showed that the profile of elemental distribution of the carbides for the untempered MZ was different from that of the tempered MZ. More uniform peak levels of carbides profile investigated in the tempered MZ (Figure 6.4-under) than that in the untempered MZ give an indication that the carbides (W, V, and Cr) distributed evenly in the structures after tempering. It also appears from the EDS results in Figure 5 that the quantity of the fine and hard carbide precipitates in the MZ was increased after tempering. This gives an indication that the fine and thin carbides diffused into the iron lattice and precipitated under secondary phase of the tempering.

85


120 Ȟm

120 Ȟm WM, 1722

120 Ȟm

Figure 6.4 SEM microstructure and profile of the elemental distribution of the untempered MZ (upper) and tempered MZ 550 oC (under). Note: power of 3000 watt, scanning speed of 200 mm/min

6.3 The Effect of Laser Treatment on the Hardness The melted sections of the T1 samples were triple tempered in order to obtain high hardness by secondary precipitation. The experimental results indicated that the average HV0.5 microhardness measured on the MZ under each laser treatment condition progressed in the same behaviour with the tempering temperature. For discussion, the microhardness behaviour of the MZ under 3000 watt laser power is presented in Figure 6.5. It appears from the results in Figure 6.5 that the highest microhardness was produced when the melted

86


Microhardness (HV0,5)

T1 sample was tempered at the temperature of 550oC. This tempering temperature was found to be the same as that used on the conventional heat treatment for the T1 bar. 950 900 850 200 mm/min 400 mm/min 600 mm/min

800 750 525

550

575

600

Tempering (oC) Figure 6.5 Microhardness behaviours with tempering temperature of the melted T1 samples at different scanning speed under 3000 watt laser power The measurement of the microhardness was made in the cross sections of the melted T1 samples starting from the surface of the MZ going down to the substrate by passing the HAZ. Distribution of the microhardness along the thickness of the melted sections under the 3000 watt laser power and 200 mm/min scanning speed with 4 mm laser spot is described in Figure 6.6. The distribution of the microhardness for the unmelted T1 (conventional hardened T1), which was made along its thickness, is also put in the figure for the comparison. The unmelted T1 shows uniform microhardness along its thickness and its average microhardness was measured to be 770 HV0.5. However, the microhardness of the laser melted T1 was found to be varied in its three zones. Microhardness measured on the MZ before tempering was around the hardness of the unmelted T1 sample. The hardness in that zone was significantly increased when it was tempered at temperature of 550 oC and reached the value of about 900 HV0.5. The increase of microhardness on

87


the MZ after tempering can be explained as a result of the formation of fine resolidified microstructure with dense distribution of very fine interdendritic carbides, the elimination of the retained austenite, and diffusion of thin carbides into the iron lattice [16]. The same phenomenon was also found during tempering the melted X165CrMoV12 steel tool [17], and tempering the melted AISI M42 [9].

Microhardness (HV0,5)

950

TZa

900

HAZ

MZ

850

TZb Substrate

800 750 Conventionally treated Laser treated without tempering Laser treated with tempering

700 650 600 550 500 0.0

1.0

2.0

3.0

4.0

Depth under surface (mm) Figure 6.6

Microhardness characteristics of the melted T1 (under the 3000 watt laser power and 200 mm/min scanning speed) and of the conventionally hardened T1

It also appears from the results in Figure 6.6 that the microhardness of the melted T1 with tempering was decreased in the HAZ and was finally the same in behaviour as the untempered melted T1 after the depth of about 1.7 mm. The microhardness was decreased to the lowest value in the TZb (depth of about 2.3 mm), and was increased again in the substrate to reach the value of the unmelted T1 sample. The microhardness in the TZb was found to be the same as the microhardness of the annealed T1 tool steel. The drop of microhardness in the depth between 1.7 mm to 2.6 mm is the results of coalescence of carbides formed previously during tempering in its conventional heat treatment.

88


The results of microhardness measurements for all laser treatment conditions are presented in Figure 6.7. The values of microhardness showed in the figure are based on the average of 10 points of measurement performed on the MZ after three times tempering one hour each at 550oC. Figure 6.7 gives an indication that microhardness tends to slightly decrease with increasing the scanning speed. The decrease is more significant in the lower laser power compared to in the higher laser power. This phenomenon probably involves the length of interaction time during the dissolving. The lower the scanning speed, the longer is the interaction time for dissolving the primary carbides to enrich the alloying elements in the solution before solidification. However, all laser treatment conditions performed in this experiment followed by three time tempering of one hour each at 550oC provided larger microhardness compared to the unmelted T1.

Figure 6.7 Behaviour of the melted T1 samples with scanning speed after three times of one hour each tempering at 550oC under different laser power It also appears from the results in Figure 6.7 that the laser power did not significantly influence the microhardness of the melted T1 under each scanning speed performed. This phenomenon can be explained that the hardness of melted zone depends largely on the quantity and size of the primary carbides dissolved during laser melting. As the amount and size of the primary carbides 89


dissolved on the surface of the T1 samples are relatively the same, the hardness produced would be almost the same. The laser power of 3000, 4000, and 5000 watt was proved to completely melt the carbides, however their melted volumes were significantly different. This result can give a good indication for selecting the treatment conditions on which laser power and the scanning speed should be applied. Beside the microhardness, the melted volume and the cracks were also considered in determining the optimum laser treatment condition. The results of this experiment give an indication that the higher the laser power and the lower the scanning speed, the higher is the risk to obtain a defect of crack and pore [16].

6.4 Wear Resistance of the Laser Melted T1 Cutting Tool Considering the results on the microhardness, melt volume and defects produced by each laser treatment condition applied on the surface of T1 samples, the combination of 3000 watt laser power and 200 mm/min scanning speed followed by three time tempering of one hour each at 550oC was considered to be a good condition. This combination provided microhardness of MZ of 900 HV0.5, depth of MZ of about 1.2 mm without any defects. Therefore, that combination was applied for melting the T1 tool followed by tempering. The results of wear test of the unmelted T1 (conventional heat treated T1) cutting tool and the tempered melted T1 cutting tool in peeling the Beech wood are presented in Figure 6.8. It appears that the tempered melted T1 peeling tool was smaller in clearance wear and suffered less fracture of cutting edges compared to the unmelted T1 peeling tool (Figure 6.9) [18].

90


Figure 6.8 Clearance wear behaviours of the tools tested with cutting length The wear mechanism of the tempered melted T1 peeling tool was caused by mechanical abrasion. The worn edges of the tempered melted T1 revealed a slight rounding edge with smooth surfaces indicating the action of abrasion without any fracture failures. Different wear mechanism was observed in the unmelted T1 peeling tool. The wear of the unmelted T1 was determined by abrasion with many fractures. At the end of cutting test (length of cut of 2 km) the worn edges of the unmelted T1 was corrugated indicating many fractures. The possible cause of this fracture was due to hard knots present in the Beech wood. The higher in hardness of the tempered melted T1 compared to the unmelted T1 caused its higher resistance to retain abrasion. The higher resistance to fractures of the tempered melted T1 compared to the unmelted T1 was considered due to its homogeneous and extremely fine microstructures with less porosity. The perfect microstructures of the tempered melted T1 would provide high toughness and strength, which lead to the elimination of the fracture wear. It was noted in another study that high-alloy martensite, fine austenite grain size, and finely dispersed carbides produced in the melted zone will contribute to high hardness, better toughness and low coefficient of friction of the melted zone which lead to its high wear resistance [11, 19].

91


Cutting Cutting Edge

EdgeEdge Chipping

Abrasion

Figure 6.9 Wear patterns of the tools tested before cutting (left) and after 2 km of cutting (right) for conventional T1 (upper) and for melted T1 (below)

References 1. Davis JR (1998) Tool Material’, 2nd edn, USA, ASM International. 2. Darmawan W, Tanaka C, Usuki H, Ohtani T (2001) Journal of Wood Science. 47: 94–101. 3. Sheikh-Ahmad JY, Stewart J S (1995) Proc. 12th Int. on Wood Machining Seminar, Kyoto, Japan, October 1995, Kyoto University, Paper 282– 291. 4. Usenius S, Korhonen AS, Sulonen MS (1987) Surface and Coatings Technology, 33: 140–143. 5. Hecht J (1999) The Laser guidebook’, 2nd edn, USA. McGraw-Hill. 6. Dahotre NB (1998) Lasers in Surface Engineering, 1nd edn, Canada, ASM International. 7. Kac S, Kusinski J (2003) SEM and TEM microstructural investigation of high-speed tool steel after laser melting’, Report 9822, Krakow-Krynica, Poland. 8. Quesada J, Monthavon G, Cornet A, Freneaux O, Jacura O, Blanc M (1989) Laser surface treatments of tool steel’, Report 45, ENSAM, Cluny, France.

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9. R. Colaço, E. Gordo, E.M. Ruiz-Navas, M. Otasevic and R. Vilar: Wear, 260: 949–956. 10. Kusinski J (1995) Applied surface science, 86: 317–322. 11. Hsu M, Molian PA (1988) Wear, 127: 251–254. 12. Singh R, Kurella A, Dahotre NB (2006) Journal of Biomaterials Applications, 21: 49–73. 13. Bendikienė R, Chodočinskas S, Pupelis E (2004) Proceedings of the Estonian Academy of Sciences Engineering, 10: 3–9. 14. Gong-ying L, Wong TT, Gen A (2004) Transactions of Nonferrous Metals Society of China, 14: 885–889. 15. Gadag SP, Srinivasan MN, Mordike BL (1995) Materials Science and Engineering A, 196: 145–151. 16. Darmawan W, Quesada J, Marchal R, Usuki H (2007) Characteristics of Laser-Melted AISI-T1 High Speed Steel and its Wear Resistance in Peeling Wood. Surface Engineering. Vol. 23 (2): 112–119. 17. Wu R, Xie C, Cai W (2000) Materials Science and Engineering A, 278: 1–4. 18. Darmawan W, Quesada J, Rossi F, Marchal R, Machi F, Usuki H (2009) Performance of Laser Treated AISI-M2 Cutting Tool for Peeling Beech Wood. Holz als Roh- und Werkstoff. Vol 67 (3): 247–255. 19. Colaço R, Vilar R (2005) Wear, 258: 225–231.

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VII. Characteristics of LaserCladding M2 High Speed Steel Cutting Tools For many wood machining processes, the interest of tool steels remains very important because of their good tool edge accuracy and easy grinding. High speed steel (HSS) is still manufactured for cutting tools in the woodworking industry. HSS is a common cutting tool material for planer blades, moulder and shaper knives, router bits, and rotary veneer knife. The interest of HSS for wood cutting tools remains very important because of their good tool edge accuracy and easy grinding. The main problem is their low resistance to both mechanical and chemical wearing. Resistance of HSS cutting tools to wearing is a primary concern in the applicability of the HSS cutting tools to a wood cutting operation. Therefore, the chemical stability of the HSS cutting tools, as well as its hardness and toughness, must be considered. Increasing their performance, a laser melting and cladding applied on the tool edges was discussed in this chapter. Surface cladding is now being a viable alternative to improve the quality of the surface properties. This cladding process typically involves the use of powder that is efficiently melted by a laser beam and precisely deposited onto the substrate material to form a well-bound clad layer. The microstructure of the clad depends on the alloy that is chosen to form it. Clad layer samples produced using a diode laser have an entirely fine dendritic microstructure [1]. In another report, Navas et al. [2] stated that microstructure of AISI M2 clad was formed by a cellular-dendritic zone composed of fine dendrites transformed to martensite and carbide, and AISI 431 clad presented a microstructure formed mainly by martensite with retained austenite. The laser AISI-M2 clad showed greater wear resistance than the AISI 431 clad due to its dense network of carbides, which effectively support the load applied during the test. In the AISI M2 clad, plastic deformation was the dominant wear mechanism, and oxidation the secondary mechanism. Otherwise the mechanism for the


samples cladded with AISI 431 was predominantly plastic deformation with a high contribution of abrasion. The friction coefficient of both layers remained approximately constant during the test, not experiencing significant changes with the load. McCay et al. [3] detailed that characteristic of laser cladding processes include low thermal distortion, minimal metallurgical degradation to the substrate, relatively high deposition rates, and rapid solidification rates associated with the deposit. These combinations of attributes make the laser cladding process an ideal candidate for a wide variety of applications. However, the ability to obtain the demanding properties required for wood cutting tool purposes has provided a formidable challenge. Initial research was conducted in conjunction with the IREPA Laser, Strasbourg, France. AISI-M2, which is widely used for wood machining purposes in form of drills, cutters and end mills, was a tool material selected for the research Tabel 7.1. An annealed M2 bar was melted and M2 powders were cladded onto the surface of low carbon steel substrate (AISI L2) by using a Diode Laser Beam. A combination of parameters (laser power, scanning speed, diameter of laser spot, and powder feed rate), which is considered to be importance in affecting the performance of the laser treatment, was fixed after some preliminary test. Table 7.1 Specifications of AISI-M2 high speed steel for investigation Steel materials for investigation Hardened M2 bar (annealed at 900oC, austenitized at 1220oC followed by three time of one hour each tempering at 560oC) M2 powder L2 steel bar

96

Chemical Composition (%wt) Fe=Bal, C=0.88, Si=0.25, Mn=0.30, P=0.02, S=0.001, Cr=4.04, W=6.13, Mo=4.92, V=1.85 Fe=Bal, C=0.85, Si=0.20, Mn=0.30, Cr=4.35, W=6.30, Mo=5.00, V=1.90 Fe=Bal, C=0.45, Si=0.35, Mn=0.15, Cr=0.80, Mo=0.20, V=0.15

Dimension or Size 500 x 40 x 10 mm

53–150 µm 500 x 40 x 10 mm

VII. Characteristics of Laser-Cladding M2 High Speed Steel Cutting Tools

M2 powders, whose composition is indicated in Table 7.1, were used to build multiple clad layers on the groove of the substrate. Numerical controlled coaxial laser cladding system was applied in the experiment. The laser used in the study [4] was a continuous wave diode type with an output power up to 3 kW. Its configuration consisted of two diode stacks with a wavelength of 808 and 940 nm. The laser beam was focused 15 mm above the substrate resulting in laser spot diameter of 3.2 mm. The powder was injected into the nozzle by means of an argon gas stream, which transported it from a feeding unit. An additional stream of the same gas was flowing through the nozzle in order to avoid surface oxidation. In order to create a clad layer of 2 mm in height, the powder feed flow was adjusted to be 23 g/min, and the laser beam was set up at speed of 600 mm/min and power of 2800 watt (Table 7.2). Overlapped tracks were made by successive scans until the groove was covered by the M2 clad layer. Table 7.2 Laser melting and cladding conditions Treatment Melting Cladding Laser power (watt) 2800 2800 Scanning speed (mm/min) 200 600 Laser spot diameter (mm) 4 3.2 Powder feeding rate (g/min) 23 Width of clad overlap (mm) 2.2 Tempering after treatment (oC) Three times one hour Three times one hour each at 560 each at 560

M2 peeling tools with 20o sharpness angle and 0o clearance angle were prepared from a M2 melted sample, M2 clad sample, and M2 conventionally hardened bar by cutting and grinding technique. Both the M2 melted and M2 clad peeling tools were tempered at 560 oC for three time one hour each before wear test. The wear resistance tests for the M2 peeling tools were performed in a peeling microlathe (Figure 7.1). Wood disks in diameter of 380 mm were peeled to produce veneer of 0.5 mm in thickness at a linear speed of 1 m/s and feed of 0,1 mm/rev (Table 7.3). At every specified cutting length (200 m) the peeling was stopped and the amount of wear on the clearance face of the peeling tools was measured under an optical video microscope. The illustration of wear measurement on the clearance face of the cutting tools was depicted in the previous paper [5]. Their worn edges were also investigated under SEM

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and optical profilometer to characterize their wear patterns. The peelings of the disks were continued up to total cutting length of about 2 km. Table 7.3 Woods specifications and peeling conditions Wood specification Wood species Beech Density (g/cm3) 0.71 Moisture content (%) 52 Thickness of disk (mm) 18 Diameter of disk (mm) 380

a

Peeling condition Cutting speed (m/s) 1 Feed speed (mm/min) 0.1 Thickness of cut (mm) 0.5 Cutting geometry Sharpness angle 20o Clearance angle 1o

b

Figure 7.1 Schematic diagram of the peeling test on a microlathe At every 200 m of the cutting length, X axis force component (parallel force) was also recorded. A laser profilometer measurement system (WYKO NT1100) was used to measure both the roughness on the face of veneer and the roughness on the cutting edges of the tool.

7.1 Microstructures of the Laser Treated M2 SEM microstructures of the M2 conventional, M2 melted and M2 Clad are shown in Figure 7.2. Metallurgical structure obtained by conventional heat treatment is a ferritic polycrystalline with coarse primary carbides (Figure 7.2a). These carbide grains were varied in size and shape, and were uneven in distribution. The result in Figure 7.2b shows the microstructure of the

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M2 melted, in which whole primary carbides were completely dissolved during laser melting. The microstructure of the M2 melted was observed to reveal fine iron dendritic structure. Almost the same microstructure was also observed in the M2 clad, in which powder metals were completely melted to form a clad layer on the surface of the substrate presenting fine iron dendritic structure (Figure 7.2c). Fine carbide networks formed due to dissolution of the primary carbide grains during laser melting were also observed in the melted and clad microstructure (Figure 7.2b–c). Investigation of the melted and clad microstructure under SEM-EDS reveals that the carbide networks (interdendritic zone) presented lamellar eutectic structure composed of Fe3W3C carbides [4]. Mainly the laser applications directed towards improving the surface properties of materials involve surface alloying and melting [6].

Interdendritic zone

a

b

c

Figure 7.2 SEM microstructures of the M2 conventional (a), M2 melted (b), and M2 clad (c)

7.2 Microhardness of the Laser Treated M2 An M2 melted and M2 clad sections were triple tempered in order to obtain higher hardness by secondary precipitation. The microhardness behaviours of the M2 melted, M2 clad and M2 conventional are presented in Figure 7.3. Measurements of the microhardness were made on the melted or clad zone along the thickness at a cross section of the tempered section and the results were in Figure 7.3a. Measurements were also made on the melted and clad zone along the surface parallel to laser scanning direction, and the results were in Figure 7.3b. It appears from the results in Figure 7.3 that the microhardness of the tools fluctuated slightly, however the amount of hardness both along its thickness and its length surfaces was almost the same. Average microhardness of the M2 conventional was measured to be 815 HV0.5. The average microhardness of the M2 melted and M2 clad was found to be the

99


same of about 840 HV0.5. The increase of microhardness for the M2 melted and clad can be explained as a result of the formation of fine resolidified microstructure with dense distribution of very fine interdendritic carbides, the elimination of the retained austenite, and diffusion of thin carbides into the iron lattice [4,7].

Microhardness HV0.5

850

M2-conv M2-melted M2-clad

860

a

840 830 820 810 800 790

Microhardness HV0.5

860

850

b

840 830 820 810 800

M2-conv

M2-melted

M2-clad

790

1 2 3 4 5 6 7 8 9 10 11 12 13 14

Point of Measurement

1 2 3 4 5 6 7 8 9 10 11 12 13 14

Point of measurement

Figure 7.3 Microhardness behaviours of the M2 laser treated and M2 conventional along the thickness (a) and along the length (b) of the sections

7.3 Wear Resistance of the Laser Treated M2 Cutting Tool The result of wear test for the cutting tools tested in peeling the Beech wood is presented in Figure 7.4. The results in Figure 7.4 indicate that the amount of clearance wear increased with increasing in cutting length. The M2 melted and M2 clad provided better performance especially in reducing the progression of clearance wear than the M2 conventional in peeling the Beach wood [4,7]. It was noted by Hsu and Molian [8] that the life of a laser melted AISI-M35 bit for cutting a steel material had to be 20 to 125% higher than if conventionally hardened.

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Figure 7.4 Clearance wear behaviours of the tools tested with cutting length Though the M2 conventional showed almost the same wear progress near beginning of the cutting, however, the clearance wear of the M2 conventional increased markedly and exceeded the clearance wear of the M2 melted and M2 clad, which retained gradual wear progresses, after 1 km of cutting length. The M2 conventional tool suffered clearance wear of about 610 µm at the 2 km cutting length, otherwise, the clearance wears were about 500 µm and 490 µm at the 2 km cutting length for the M2 melted and M2 clad respectively. The higher in hardness and better in microstructures of the M2 melted and M2 clad compared to the M2 conventional would be the reason for this phenomenon.

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a1

1 mm

b1

1 mm

c1

1 mm

1 mm

a3

50 Pm

b2

1 mm

b3

50 Pm

c2

1 mm

c3

50 Pm

a2

Figure 7.5 Photo micrograph of worn cutting edges before cutting (a1,b1,c1 and after 2 km of cutting length (a2,b2,c2), and the SEM micrograph of the worn edges at the 2 km cutting length (a3,b3,c3). Note: a=M2 conventional, b=M2 melted, c=M2 clad The results of wear patters in Figure 7.5 shows that the M2 melted and M2 clad cutting tool suffered less fracture of cutting edges compared to the M2 conventional cutting tool. The wear mechanism of the M2 melted and M2 clad cutting tool was caused by mechanical abrasion. The worn edges of the M2 melted and M2 clad (dark colour along the cutting edge) revealed a slight rounding edge with smooth surfaces indicating the action of abrasion without any fracture failures (Figure 7.5b2–c2). Different wear mechanism was observed in the M2 conventional peeling tool. The wear of the M2 conventional was determined by abrasion with many fractures at the end of cutting test (Figure 7.5a2). SEM micrograph for the worn edge of the M2 conventional (Figure 7.5a3) shows that the cutting edge was corrugated indicating many fractures. Its susceptible to fracture was considered due to the fact that any carbide grains on the edge were detached by mechanical abrasion during peeling (Figure 7.5a3). A possible cause of the detached grains was due to abrasion by hard knots present in the Beech wood. The higher in hardness of the M2 melted and M2 clad compared to the M2 conventional caused their higher resistance to retain abrasion. The higher resistance to fractures of the M2

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melted and M2 clad compared to the M2 conventional was considered due to its homogeneous and extremely fine microstructures with very fine grain sizes [4,7]. Ahman [9] stated that a dendritic structure with a uniform and dense dispersion of fine carbides grains after laser melting was found to improve toughness of M2 steel. In another study, Kusinski [10] found that the surface of M2 steel after laser melting was better in erosive, fatigue resistance, and friction. It appears from the results in Figure 7.2b–c that fine, homogenous, and evenly dispersed carbides produced in the structure of M2 melted and M2 clad were considered to contribute to high hardness, better toughness and low coefficient of friction of the M2 melted and M2 clad which lead to their high resistance against abrasion and fracture.

7.4 Behaviour of Parallel Cutting Force (Vertical force) The results in Figure 7.6a show that the parallel force generated by the cutting tools increased in proportion with an increase in cutting length. The M2 conventional, compared to M2 melted and M2 cald peeling tools, tended to be larger in parallel force especially near end of peeling due to its higher clearance wear and uneven surfaces of its cutting edge because of fracture. The results in Figure 7.6b give an indication that the parallel force of the tools increased linearly with increasing in the clearance wear. The progresses in the parallel force with increasing in the clearance wear were almost the same among the tools as indicated by their same coefficients of regression. This gives an indication that the tools generated almost the same parallel force as long as their amount of clearance wear are same. Though considerable differences in parallel force among the tools were not observed in peeling the Beech wood, however their significant increase in the parallel force with increasing in cutting length or amount of clearance wear could be a good indicator for monitoring the progress of edge wear during peeling the Beech wood.

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Figure 7.6 Behaviour of the parallel force with cutting length (a) and clearance wear (b) for the M2 laser treated and M2 conventional cutting tools in peeling Beech. Note: y=parallel force, x= clearance wear, r=correlation coefficient [4]

7.5 Roughness (Ra) Value of the Veneer Surfaces Producing proper surface finish is an important part of the wood machining process. The surface finish of veneer is closely related to wood adhesion in production of plywood. The final surface roughness of veneer is considered as the sum of independent effects of geometry of tool, linear speed and feed rate, and wood characteristics. Factors such as linear speed, feed rate, and depth of cut (veneer thickness) that control the cutting operation could be set-up in this experiment. However, factors such as tool wear, the material properties of both tool and workpiece are uncontrolled. It appears from the result in Figure 7.7a that the roughness of veneer (Ra) increased with increasing in the length of cut. In the beginning of cutting when the cutting tools are same in sharpness, the roughness of veneer produced are almost same. However, after cutting length of 800 m the M2 conventional tended to produce veneer with rougher surface compared to the M2 melted and M2 clad, which produced almost the same roughness of veneer. This phenomenon was caused by fact that the M2 conventional suffered higher clearance wear of clearance wear and edge fracture than the M2 melted and M2 clad. Apparently, the roughness of the cutting edge of the tools also toke an important role in determining the roughness of veneer produced (Fig7.7b). The roughness of veneer increased

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in prpoportion with an increase in the roughness of the cutting edge. Before cutting, M2 conventional presented slightly rougher cutting edge then the M2 melted and M2 clad. This could be caused by the presence of carbide grains in the edge, which were not completely cut by the grinding machine. The difference in edge roughness between the M2 conventional and the M2 laser treated became larger as the cutting length was increased. At the 2 km cutting length, the edge roughness of the M2 conventional was twice larger than that of the M2 laser treated. M2 conventional, which suffered the higher roughness of cutting edges produced rougher of veneer surfaces. The corrugated edge of the M2 conventional due to fractures was considered to crush grains on the veneer surfaces, which resulted in the increase in the roughness [4].

22

25

a

19 16


13 10

0.2 0.4 0.6 0.8 1.0 1.2 1.4 1.6 1.8 2.0

Length of cut (km)

Ra of veneer (µm)

Ra of veneer (µm)

25

b

22 19


16 13 10

0

10

20

30

Ra of cutting edge (µm)

Figure 7.7 Behaviour of the Ra of veneer (a) and Ra of cutting edges (b) of the M2 cutting tools tested in peeling Beech wood

References 1. Pinkerton AJ, Li L (2004) Multiple-layer cladding of stainless steel using a high-powered diode laser (An experimental investigation of the process characteristics and material properties. Thin Solid Film. 454: 471–476. 2. Navas C, Conde A, Fernandes BJ, Zubiri F, Damborenea JD (2005) Laser coating to improve wear resistance of mould steel. Surf. Coat. Technol. 194: 136–142. 3. McCay MH, Dahotre NB, Hopkins JA, McCay TD, Riley MA (1999) The Influence of Metals and Carbides During Laser Modification of Low Alloy Steel. Journal of Materials Science. 34: 5789. 105


4. Darmawan W, Quesada J, Rossi F, Marchal R, Machi F, Usuki H (2009) Performance of Laser Treated AISI-M2 Cutting Tool for Peeling Beech Wood. Holz als Roh- und Werkstoff. Vol 67 (3): 247–255. 5. Darmawan W, Usuki H, Marchal R, Quesada J (2008) Clearance Wear and Normal Force of TiN-coated P30 in cutting hardboard and woodchip cement board. Holz als Roh- und Werkstoff . Vol 66 (2): 89–97. 6. Kac S, Kusinski J (2003) SEM and TEM microstructural investigation of high-speed tool steel after laser melting. Materials Chemistry and Physics. 81: 510–512. 7. Darmawan W, Quesada J, Rossi F, Marchal R, Machi F, Usuki H (2009) Improvement in Wear Resistance of AISI-M2 Cutting Tool by Laser Treated. Journal of Laser Technology. Vol 21 (4): 176–182. 8. Hsu M, Molian PA (1988) Tool life of conventionally heat treated and laser melted of M35 bit. Wear. 127: 251–254. 9. Ahman L (1984) Microstructure and its effect on toughness and wear resistance of laser surface melted and post heat treated high speed steel. Metallurg Trans A. 15(10):1829–1835 10. Kusinski J (1995) Microstructural, chemical composition, and properties of the surface layer of M2 steel after laser melting under different conditions. Applied surface science. 86: 317–322.

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VIII. New Design of Helical Edge Milling Cutters for Planing Wood Optimisation and improvements in wood cutting operations are characterized by various approaches, including design of cutting tool, selection of the cutting tool material, and application of machining conditions. All of the approaches lead to much higher productivity, more economical cutting and reduction of the overall machining cost, in which resulted from better efficiency, stability, accuracy, and tool life during the cutting processes. For economical and high performance cutting, all parts involved in the cutting processes should be selected and optimised. Among the parts, design of cutting tool edge involved in the cutting processes would be very important. Today conventional designs of peripheral milling cutters with two or more straight cutting-edges are widely used in the wood working industry for planing purposes. The manner of contact between this straight configuration of cutting edge and the work piece is a piecewise continuous curve. The straight cutting edge hits and intermittently engages the surface of the work piece during planing. As the results, the noise generated during the planing process tends to be very loud and the cutting force generated is relatively high [1]. This straight configuration leads to machined surface quality problem due to high splitting, compressing and damaging the wood cell structure near the surfaces. Costly sanding procedures are needed. Peripheral milling principle of the straight configuration leads to produce extreme flight-speeds of the formed chips by tangential acceleration. Suction system is mostly unable to catch the chips completely. Dust emission will occur and high energy is necessary to increase the efficiency of the process. In addition each knife-edge is under sudden hit of possible wood knots and compression of cellular surface layers, which lead to early damage and dull cutting tool edges. Minimizing this edge problem phenomena requires a lot of sharpening processes and tool adjustment to improve the working accuracy.


Consequently a lot of efforts in minimizing the problems of the straight configuration should be needed. One approach dealing with a new design of helical edge has been being developed. Some research works and investigations were done to find out the effect of inclination angles of the helical cutting tool edge for wood cutting applications. Mostly the research works were focused on energy behaviour, surface quality, dust emissions, and noise emissions with respect to the varied inclination angles. The first general overview was investigation of the helical edge design with inclination angle up to 30° for wood chipping application [2,3]. They noted that an increase in the inclination angle leads to an increase in the passive force (axial direction), however the vibration was reduced and the noise level was lowered. It was reported in another study that noise reduction of more than 10 dB(A) was observed when cutting wood with helical edge milling tools of 18° inclination angle [4]. Starting a few years ago, research activities on the investigation of dust, chip, noise and force behaviours in milling operation using helical edge with inclination angles between 0° to 45° were reported [5,6]. It was noted in these studies that inclination angles between 5° and 10° were considered to be useful in lowering the dust emissions. Inclination angles larger than 10° lead to chip compression resulting high axial forces on the cutting tool. The increase in the inclination angle from 0° to 8° leads to a significant decrease of noise emissions. Recently some publications have known to report the results on the influence of the inclination angle to performance of cutting processes. It was found that helical edge milling tool with inclination angle of 0° to 8° did not provide a significant difference in the cutting energy and the chips morphology when milling with shank tools of small diameter [7]. It was reported in another study that helical edge design of a milling tool was reported to provide better surface quality compared to straight cutting edge [8]. In milling wood against the grain, the greater the helical angle the smoother was the machined surface. It could be considered that when using cutting edges with the helical configuration, the cutting edge penetrates gradually into the work piece with a step-wise force increase, reaching a maximum value that will be lower than that achieved with a straight cutting edge. When cutting edges engage the surface of the work piece gradually, the resultant cutting forces will be lower, the tools will be always under contact, stability will be improved, vibration

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VIII. New Design of Helical Edge Milling Cutters for Planing Wood

will be reduced and the required machine power during the milling operation will be lowered. Because of the use of the unique geometry of cutting edge, the wood chips will be easily deformed and the resulting machined surfaces will be flat and smooth. The effect of helix angle on cutting forces and consequently on tool life for wood machining in previous studies was decided to deepen the analysis because tool makers need to understand such effects [9]. The oblique cutting edge is divided into small differential segments. For each oblique element, cutting forces are formulated to introduce the density variation of wood based materials: cutting constants, obtained by measuring cutting forces for various orthogonal millings on MDF and particleboard, are written in function of height coordinate of the board thickness. To validate the three orthogonal cutting forces calculated from measured constants, a comparison with new experimental values is necessary. The predictions obtained indicate the effects of helix angle on the tool life. The effect of the helix angle of a router bit on chip formation and electric energy per volume (specific energy) under different feed speeds and cutting depths during the milling of maple and China fir by a computer numerically controlled (CNC) router was investigated [10]. The peripheral cutting edge of router bits were custom-made at helix angles of 0°, 2°, 4°, 6°, and 8°. The feed speed varied from 600 to 4800 mm/min, and the depths of cut were set at 1, 2, 3, 4, and 6 mm. The chips were classified by sieving into a flake type, a splinter type (5 and 10 mesh), a flow type (20 and 40 mesh), and a granule type (< 40 mesh). As the feed speed and the cutting depth increased for the five router bits, more chips of the flake type and the splinter type were produced. However, the number of granule-type chips under the larger helix angle was reduced. The energy per volume removed (specific energy) increased with the feed speed and the depth of cut while milling maple and China fir. More specific energy per cubic centimeter was consumed under the lower feed speeds and the smaller depth of cut. The specific energy can be expressed as a negative power function of the feed speed or the cutting depth for maple and China fir.

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Pay close attention to the above phenomenons, and to the fact that the previous research works were done limited to the inclination angles between 0° to 45°, therefore a new design of helical edge of milling cutter with extreme inclination angles (45° to 85°) has been developed at the Technische Universität Dresden (Dresden University of Technology) [11,12] and their performance was tested and discussed in this chapter. The scope of this chapter is to discuss the effect of inclination angles on the chip flow, cutting power, noise emission, edge wear and surface roughness characteristics of the extreme helical edge of the milling cutter in planing wood [13]. The milling cutters consisted of one solid cutting edge with inclination angles of 0° (conventional edge), 45°, 55°, 65°, 75°, and 85° (Figure 8.1). Other geometries of the cutter heads are shown in Table 8.1. The wood species machined was spruce (Picea abies) of 12 % in moisture content. An up-milling process was performed by setting the rotation of a moulder spindle in clockwise direction and by feeding the lumber samples in the opposite direction with rotation of the spindle. The lumber samples were planed along the length on their side surfaces. Schematic diagram of the milling test is depicted in Figure 8.1, and conditions of the milling are shown in Table 8.2. Table 8.1 Specifications of conventional and helical edge milling cutters for investigation Milling tool material Hardness Cutting circle diameter d Number of cutting edge z Width of milling tool bmax Geometry of the edges Inclination angle Orthogonal rake angle Orthogonal clearance angle

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Cold steel X155CrMo12-1 60 HRC 125 mm 1 75 mm 0° (conventional edge), 45°, 55°, 65°, 75°, 85° 27° 13°


150 mm

Cutting edge

10 mm

a

1 mm

0°

45°

65°

2000 mm

55°

75°

85°

b

Figure 8.1 Schematic diagram of the up-milling process (a) and milling cutters for investigation (b) Table 8.2 Milling test conditions Adjusted parameters Cutting speed vc (m/s) Feed speed vf (mm/min) Depth of milling ae (mm) Width of milling b (mm) Inclination angle (°)

Conditions for noise and power test wear test 39 39 4, 10, 16 6 1 1 10 10 0, 45, 55, 65,75,85 0, 65, 75, 85

8.1 Chip Flow Digital video images analysis indicated that the milling cutters tested produced different behaviours of chip flow. The chips flow in different direction, angle, and speed during the cutting process, as illustrated in Figure 8.2. When the

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lumber samples had been planed using the conventional edge, the chips flow in tangential direction with larger area of flow (sector area between the two arrows), and with higher speed of flow. Consequently the chips tended to scatter around the cutting point, which in turn could affect the performance of rest cutting processes. The schematic redrawing results of video images in Figure 8.2 show that the sector area of the chip flow became smaller and the investigated speed of chip flow was decreased, as the edge inclination angles were larger. The smallest area of flow and speed of flow with nearly axial direction were generated by the milling cutter of 85° inclination angle. This result gives an indication that the kinetic energy of the chip flow would be more lower with increasing in the edge inclination angle. It was also observed that the chips moved in a regular parabolic way during cutting tests using milling cutters with extreme edge inclination angles (65°, 75°, and 85°). The chips were collectively deposited in a place at a distance of about 30 cm to 50 cm away of the cutting point. It could be considered that a complete capture of dusts and chips with enormous saving of energy consumption for the required suction system would be realized, as a proper hood would be placed around the sector of expected chip flow [13]. The chip flow velocity decreased rapidly when increasing the edge inclination of the tool from 37,5 m/s (λS = 0°) to approximately 16 m/s (λS = 85°). The analysis on the particle sizes generated during the cutting test showed a content of less than 0,3 % of dangerous particles smaller than 0,1 mm. All in all the size distribution of the large particles fractions depend strongly on the cutting direction and conditions as well as the inclination angle too. For deeper quantification further investigations are necessary. Fortunately a serious problem in the chips suction system and a high suction energy required during milling with a conventional cutter could be improved by using milling cutters with extreme edge inclination angles.

0°

65°

75°

85°

Figure 8.2 Schematic redrawing of the measured chip flow sector from images during milling with different too 112


8.2 Cutting Power Consumption The results in Figure 8.3a show that helical edge of milling cutters with inclination angle between 45° to 65° generated almost the same amount of cutting power as the conventional edge of milling cutter. However the cutting power consumption increased gradually for the inclination angle of 75° and 85°. The helical edge with 85° inclination angle generated cutting power consumption twice larger than the others at the same feed speed. This fact is considered due to the cutting edge with most extreme inclination angle (85°) being higher in forces caused by the increase in frictions between the chips and the gullet surface. The change in cutting edge engagement from parallel the grain to the nearly perpendicular the grain could be also responsible for the higher power consumption for the milling cutter of 85° inclination angle. It also appears from the result in Figure 8.3a that cutting powers generated by the extreme inclination angles (65°, 75°, and 85°) were observed to raise by increasing in the feed speed. This fact is considered to be due to the effect of higher chip thickness machined by the cutting edge as the feed speed was increased. However, the energy per volume of chips removed (specific energy) for the milling cutters of both the conventional and the helical edges was lowered when the feed speed was increased (Figure 8.3b). These results indicate that planing the wood at higher feed speed would be more energyeconomical because of lower specific energy consumed [13]. The specific energy value of the tool with 55° inclination resulted a difference to the general tendency at feed speed of 4 m/min (Figure 8.3b). This can only be explained by some unintentional changes of edge or tool properties or measurement variance during the tests.

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Cutting Power (W)

600 500 400

4 m/min 10 m/min

a

16 m/min

300 200 100 0 0 5 1015 20 25 3035 40 45 5055 60 65 7075 80 85

Inclination Angle (°)

Specific Energy (Ws/cm³)


60 50 40

4 m/min 10 m/min

b

16 m/min

30 20 10 0 0 5 10 15 20 25 30 35 40 45 50 55 60 65 70 75 80 85

Inclination Angle (°)

Figure 8.3 Cutting power (a) and specific energy (b) of the milling cutters without and with an increase in inclination angles at different feed speeds

8.3 Noise Emissions The results in Figure 8.4 indicates that the conventional milling cutter generated higher noise levels compared to helical milling cutters during the cutting test. It was measured that the noise levels were decreased when the inclination angle for the helical milling cutters was increased. It appears from the results in Figure 8.4 that milling cutter with inclination angle of 75° generated a minimum noise level. However, the noise level was found to slightly increase for the 85° inclination angle. This could be caused by increased passive forces which initiate work piece vibrations perpendicularly to the feeding direction within the working plane. The higher noise level generated by conventional milling cutter is caused by intermittent hammering and hit of its straight edge into the surface of the work piece. Otherwise the helical edge of milling cutter engaged and penetrated gradually into the surface of the work piece, which causes the noise to be going down to a lower level [13]. The differences in noise pressure level between conventional edge and helical edges were up to 10 dB(A). This result indicates a significant reduction in noise because a difference of noise pressure level by 6 dB(A) means a doubling of loudness according to the logarithmic definition of noise level.

114


Noise Level (dB(A))

106 104 102 100 98

4 m/min

96

10 m/min 16 m/min

94 92

0 5 10 15 20 25 30 35 40 45 50 55 60 65 70 75 80 85

Inclination Angle (°) Figure 8.4 Noise pressure level of the milling cutters without and with an increase in inclination angles at different feed speeds It also appears from the results in Figure 8.4 that the milling cutters generated lower noise levels during cutting at the lower feed speed compared to the higher feed speed. The differences of around 5 dB(A) were observed between the feed speed of 4 m/min and 16 m/min. The high noise levels generated during high feed speed cutting are caused by large impact forces being imposed on the milling cutters for the high feed speed cutting. The same phenomenon was also noted by Darmawan and Tanaka [14] in reporting the discrimination of coated carbide tools wear by the features extracted from parallel force and noise level.

8.4 Wear Resistance The results in Figure 8.5 indicate that the tool edge radius increased with increasing in cutting length. The helical edge milling cutters provided better performance especially in reducing the progression of edge radius than the conventional milling cutter in cutting the spruce wood. The conventional milling cutter attained edge radius of about 40 µm in cutting spruce at the 1280 m cutting length. Though the helical edge milling cutters showed almost same edge radius wear progress, however there was a tendency that a larger inclination angle provided better edge radius resistance than the smaller inclination angle [13]. It appears that the edge radius of the helical edge of 85°

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inclination angle was slightly lower than that of the other helical edges (Figure 8.5). The helical edge of 85° inclination angle attained edge radius of 26 µm at the 1280 m cutting length. Table 8.3 Linear regression equations and correlation coefficients according to Figure 8.5 (Range of analysis between cutting length of 320 and 1280 m) Milling tool Conventional edge Helical edge of 65o Helical edge of 75o Helical edge of 85o

Linear equation y = 0.0074x + 30.5 y = 0.0051x + 25.5 y = 0.0040x + 23.0 y = 0.0040x + 21.0

r 0.997 0.994 0.964 0.964

y = edge radius, x = cutting length, r = correlation coefficient

After a degressive increasing of edge radius between the cutting length of 0 m and 320 m the progress of edge wear (radius change) can be described as a linear function. In this range linear regression equations and its correlation coefficients according to Figure 8.5 are summarised in Table 8.3. The results show that the regression coefficients for the edge radius linear equations depicted by the milling cutters varied from 0.0040 to 0.0074. These variations indicate that the inclination angles of the milling cutters determined the rate of the increase in the edge radius of the milling cutters. It appears that the rate of the edge radius wear was decreased as the inclination angle of the milling cutters was larger (Table 8.3). The helical edge of 85° inclination angle would be more gradual in the increase of edge radius compared to the others. It could be considered as a reason for this phenomenon that during the cutting actions, a larger inclination angle of the helical edge would involve a longer cutting edge in engaging the surface of the wood. Further the gradual step wise engagements of edge involved during cutting with helical edge milling cutters could eliminate sudden impacts and decrease mechanical loads on the surface of cutting edges. This result gives an indication that the helical edge of milling cutters would be better in wear resistance and provide longer tool life compared to conventional edge in application for wood planing.

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Tool Edge Radius (µm)

40 35 30 25 20

0° inclination angle 65° inclination angle 75° inclination angle 85° inclination angle

15 10 5 0

0

200

400

600

800

1000

1200

1400

Cutting Length (m) Figure 8.5 The progress of cutting edge radius with cutting length for the milling cutters tested

8.5 Roughness (Ra) Value of the Planed Lumber Producing proper surface finish is an important part of the wood machining processes. The surface finish of lumber is closely related to performance of finishing in the production of furniture. The final surface roughness of lumber is considered as the sum of independent effects of geometry of tool, linear speed and feed rate, and wood characteristics. Factors such as tool edge geometry, linear speed, feed speed, and depth of cut (planing thickness) that control the cutting operation could be set-up in this experiment. It appears from the result in Figure 7.6a that the roughness of lumber (Ra) increased with increasing in the cutting length. In the beginning of cutting when the cutting tools are same in sharpness, the roughness of surface produced are almost same. However, after cutting length of 320 m the conventional edge tended to produce planed lumber with more rough surface compared to the helical edge of the milling cutters. Among the helical edge milling tools, the 85° inclination angle tended to produce the smoothes surfaces [13]. This phenomenon was caused by fact that the conventional edge suffered higher edge radius wear than the helical edge milling cutters (Figure 7.5). Apparently,

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0° inclination angle 65° inclination angle 75° inclination angle 85° inclination angle

10 9 8 7 6 5 4

0

200 400 600 800 1000 1200 1400

Cutting Length (m)

Ra of Wood Surface (µm)

Ra of Wood Surface (µm)

the roughness of the cutting edge of the tools also took an important role in determining the roughness of surface produced (Figure 7.6b). The roughness of lumber surfaces increased in proportion with an increase in the roughness of the cutting edge. The difference in edge roughness between the conventional edge and the helical edge milling tools became larger as the cutting length was increased. At the 1280 m cutting length, the conventional edge, which suffered higher roughness of cutting edges produced more rough lumber surfaces (Figure 6b). The less corrugated edges of the helical milling cutters due to gradual engagement of the cutting edges in action were considered to smoothly cut the grains on the wood surfaces, which resulted in the less roughness [13]. This result confirms with the previous report, in which the increase in the roughness of peeling cutting tools results in the increase in the roughness of wood veneers produced [15]. 0° inclination angle 65° inclination angle 75° inclination angle 85° inclination angle

10 9 8 7 20 m

6

1280 m

320 m

5 4

4

5

6

7

8

9

10

Ra of Tool Edge (µm)

Figure 8.6 The roughness of planed wood surface with cutting length (a) and with the roughness of tool edges (b)

References 1. Chen WF, Lai HY (2002) A comprehensive engineering model for the design, manufacture and assembly of helical carpenter shapers. Journal of Engineering Manufacture. 216: 1493–1504 2. Pahlitzsch G (1966) Internationaler Stand der Forschung auf dem Gebiet des Hobelns und Fräsens von Holz und Holzwerkstoffen. Holz- als Rohund Werkstoff. 24: 579–592.

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3. Pahlitzsch G, Sommer I (1966) Erzeugung von Holzschneidspänen mit einem Messerwellen-Spaner–Dritte Mitteilung: Einfluß des Neigungswinkels, des Messerschneidenwinkels und des Schnittrichtungswinkels. Holz- als Roh- und Werkstoff. 24: 158–166. 4. Heydt F, Tuffentsammer K (1979) Lärmminderung an Dickenhobelmaschinen. HK Holz- und Kunststoffverarbeitung. 5: 384– 388. 5. Heisel U, Weiss E (1989) Einfluß von Schneidengeometrie und Bearbeitungsparameter auf die Staubentwicklung bei Kehlmaschinen. HOB Die Holzbearbeitung. 12: 16–21. 6. Heisel U, Niemeyer W, Weiss E (1993) Lärm- und staubarmer Fräsprozeß mit wendelförmigen Schneiden. HOB Die Holzbearbeitung. 5: 90–98. 7. Su WC, Wang Y (2002) Effect of the helix angle of router bits on chip formation and energy consumption during milling of solid wood. J Wood Sci. 48: 126–131. 8. Cyra G, Tanaka C, Yoshinobu M, Nishino Y (1998) Effects of helical angle of router bit on acoustic emission. J Wood Sci. 44: 169–176. 9. Boucher J, Méausoone PJ, Martin P, Auchet S, Perrin L (2007) nfluence of helix angle and density variation on the cutting force in wood-based products machining. Journal of Materials Processing Technology. 189 (1–3):211–218. 10. Su WC, Wang Y (2002) Effect of the helix angle of router bits on chip formation and energy consumption during milling of solid wood. J Wood Sci. 48:126. doi:10.1007/BF00767289. 11. Fischer R, Gottlöber C, Rehm K, Rehm C (2005) A milling cutter as a screw: cutting instead of hacking. Proceedings of the 17th International Wood Machining Seminar September 26–28 Rosenheim Germany, pp 4–10. 12. Fischer R, Gottlöber C, Rehm K (2006) Schneiden statt hacken. HOB Die Holzbearbeitung. 53: 63–66. 13. Darmawan W, Gottloeber C, Oertel M, Wagenfueher A, Fischer R (2011): Performance of Helical Edge Milling Cutters in Planing Wood. Journal of Wood and Wood Products. Vol 69 (4): 565–572.

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14. Darmawan W, Tanaka C (2004) Discrimination of coated carbide tools wear by the features extracted from parallel force and noise level. Ann For Sci. 61: 731–736. 15. Darmawan W, Quesada J, Rossi F, Marchal R, Machi F, Usuki H (2009) Performance of laser-treated AISI-M2 cutting tools for peeling beech. Eur J Wood Prod. 67: 247–255. doi: 10.1007/s00107-009-0324-2.

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About the Author

WAYAN DARMAWAN Office: Department of Forest Products Faculty of Forestry Bogor Agricultural University Bogor, INDONESIA

Contact: E-mail [email protected] Phone

+622518621285

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Wayan Darmawan is a full professor at the Department of Forest Products, Faculty of Forestry, Bogor Agricultural University (IPB) since 2012. He received his doctorate at the University of Shimane, Japan in 2000. He continued his post-doctoral study at the Ecole Nasionale Superior de Art et Metier, France in 2002–2003. He was an Erasmus Mundus visiting professor fellow at Dresden University, Germany in 2009. He had been serving as Director of the Forest Products Department in 2010–2014. He has published more than 25 articles in the domains of wood machining, wood quality improvement by surface coating and two book and a book chapters. He has been a member of the SWST since 2012, published more than 25 peer review research articles.