Model of a Surface Flaw and Calculation of Topography of Its Magnetic Field for Tangential Magnetization by Alternating Magnetic Fields. Quasi-Stationary Case

A model of a surface flaw is proposed. The model describes topography of the magnetic field for tangential magnetization by an alternating magnetic field H at a frequency . It is shown that H can be represented in the form of two multipliers: one describes the dependence on the (X, Z) coordinates and the flaw parameters (the depth h and the opening width 2b) and the other describes the dependence on the electrophysical properties of the tested material in which the surface flaw is located (the specific conductivity and the relative magnetic permeability ). The computational model proposed makes it possible to describe the magnetic field topography ( = 0) reducing it to the quasi-stationary case and extend representations on the formation of magnetic fields induced by surface flaws with allowance of the magnetization frequency . The data on processing the experimental dependences in accordance with the proposed computational formulas give satisfactory results which confirm the validity of the computational model proposed in the study.     

Surface roughness measurements in finish turning

A new approach is proposed for the on-line measurement of the maximum peak-to-valley roughness,R _max, of a finished-turned surface in the feed direction. The method is based on solving the inverse problem of light scattering by using a linear least-square estimate of the angular scattered light pattern reflected from a surface. A laser system has been developed to capture the light reflected under different cutting conditions. The effects of the ambient room light as well as the workpiece's rotational speed and methods for thier compensation are also discussed. Good correlation was found between the optical and stylus-measuredR _max.Nomenclature R _max maximum peak-to-valley roughness within the sampling length - R _q RMS surface roughness within the sampling length - R _a arithmetically averaged roughness within the sampling length - _z r.m.s. surface height within the sampling length - u r.m.s. slope of the surface within the sampling length - T correlation distance of the surface, defined as the distance in which the correlation coefficient,C(), equals e^–1 - I(₁,) intensity of reflected light - I _m(₁,₂,) measured intensity of reflected light at instant - ₁ angle of incidence of laser beam - ₂ scattering angle defining a CCD pixel location (₁ and ₂ are measured with respect to the normal of the surface of the workpiece coincident with the centre of the laser beam) - v scattering vector of reflected light - _x,_z components ofv in thex andz direction, respectively - L sampling length associated with the laser spot on the surface of the workpiece - j representative location of a CCD pixel - j CCD pixel location corresponding to the mean light level - p _j density function of the light intensity of thejth pixel - wavelength of laser light - nose radius of the cutting tool - ASLP angular scattered light pattern - K correction factor for the measured light intensity - S _m standard deviation of the measured ASLP - S _c standard deviation of the ASLP calculated from an estimatedR _max - K control step size ofK - computational error, defined as =|S _m–S_c|/S _m - K _a,K_b starting and ending point, respectively, within the search range forK - K _c,K_d two points within (K _a,K_b), determined by the golden section search method - V cutting speed (m/min) - f feed rate (mm/rev) - d depth of cut (mm) - H hardness of workpiece (found on Rockwell scale C) - CCD charge-coupled device     

A generalised model of milling forces

This paper presents the development of a generalised cutting force model for both end-milling and face-milling operations. The model specifies the interaction between workpiece and multiple cutter flutes by the convolution of cutting-edge geometry function with a train of impulses having the period equivalent to tooth spacing. Meanwhile, the effect of radial and axial depths of cut are represented by the modulation of the cutting-edge geometry function with a rectangular window function. This formulation leads to the development of an expression of end/face-milling forces in explicit terms of material properties, tool geometry, cutting parameters and process configuration. The explicitness of the resulting model provides a unique alternative to other studies in the literature commonly based on numerical integrations. The closed-form nature of the cutting force expression can facilitate the planning, optimisation, monitoring, and control of milling operations with complicated tool—work interactions. Experiments were performed over various cutting conditions and results are presented, in verification of the model fidelity, in both the angle and frequency domains.Notation * convolution operator - helix angle of an end mill - _A,_R axial and radial angles of a face mill - angular position of any cutting point in the cylindrical coordinate system - unit area impulse function - ^(i–1)(–T _o) (i–1)th derivative of (–T _o) with respect to - angular position of cutter in the negative Y-direction - _L, lead and inclination angles of a face mill - angular position of any cutting point in the negative Y-direction - ₁, ₂ entry and exit angles - upper limit of cutting edge function in terms of - as defined in equation (10) - A _xk ,A _yk ,A _zk kth harmonics of cutting forces in the X-, Y-, and Z-directions - d _a,d _r axial and radial depth of cut - dA instantaneous cut area - D diameter of cutter - f _o frequency of spindle - f _t,f _r,f _a local cutting forces in the tangential, radial, and axial directions - f _x ,f _y ,f _z local cutting forces in the X-, Y-, and Z-directions - F _x ,F _y ,F _z resultant cutting forces in the angle domain in the X-, Y-, and Z-directions - F as defined in equation (5) - h derivative of height function of cutting edge with respect to - h() height function of one cutting edge with respect to - H height of any cutting point - K _r,K _a radial-to-tangential and axial-to-tangential cutting force ratios - K _t tangential cutting pressure constant - K as defined in equation (6) - p as defined in equation (6) - N number of cutting edges - r() radius function of one cutting edge with respect to - R radius of any cutting point - T cutting engagement time function of any cutting point - T _o cutting engagement time of the cutting point at =0 - T th() tooth sequence function - t _c average cut thickness - t _x feed per tooth - W _A,W _W,W _C amplitude, width and centre of a window function - W(,) unit rectangular window function - y _min,y _max minimum and maximum positions of workpiece in the Y-direction - Z _min,Z _max integration limits in the Z-direction     

Detection of the Increment of an Instantaneous Phase in a Long-Base Laser Meter of Small Vibrations

A phase meter for processing signals of a laser meter of small displacements and vibrations at long base distances is described. Vibrations of objects are transformed into small increments of a signal phase at an RF carrier, which are detected by the phase meter and are outputted as signals proportional to microvibrations in the acoustic range. At a given carrier frequency f _c = 10.7 MHz, vibrations are detected within a band f = 3 kHz. Such vibrations produce phase fluctuations of 10^–42, which correspond to magnitudes of 1 nm for a laser wavelength 10 m.     

General Shear-Thinning Dynamics of Confined Fluids

The shear properties for a number of thin fluid films under high pressure were investigated as a function of sliding velocity (shear rate) using the surface forces apparatus. It was found that the relationship between the effective viscosity _eff and shear rate in the shear-thinning regime could be expressed by a simple equation, log_10eff=C-nlog₁₀, where C4.7±0.2 and n0.9±0.1. This equation can be applied to a variety of fluid systems from simple liquids to polymer melts, which transition to glasslike phases in confined geometries. Since the effect of confinement on the slowing down of molecular motions is equivalent to that of decreasing temperature, this universally can be explained using conventional glass-transition theories for bulk fluids. Assuming the confined fluid to be in a state where dynamics are dominated by excluded volume effects, its _eff should correspond to that of the bulk at or near the glass-transition temperature. Thus, characteristic relaxation times in the system should correlate with the time scales of the primary relaxation processes associated with submolecular rearrangements, which are an essential feature of the glass transition and not very different for various fluid materials.     

Parameters for Roughness Pattern and Directionality

The average flow model is widely used on the derivations of average Reynolds type equation.There are arguments on the use of Peklenik number ( ^P ), or Bhushan number ( ^B ). In this paper, the orientation angle ( _r ) of the representative asperities (the pattern directionality) as well as the Peklenik number defined in principal directions () is utilized as the parameters to define the texture of surface roughness. An experimental procedure based on the least square method is then proposed to identify the two parameters ( _r and ). The present procedure avoids the above argument on distinguishing the isotropic asperity with the anisotropic asperity oriented with 45°. Only one additional parameter ( _r ) is needed.     

Spindle deflections in high-speed machine tools — Modelling and simulation

This research attempts to develop spindle deflection error models for high-speed machining systems. A model for determining total spindle deflection at the tool-end is presented. The model incorporates spindle bearing characteristics, shifts in ball contact angles, and centrifugal force and gyroscopic moment effects at high speeds. It uses the transfer matrix method to determine the total deflections at the tool-end based upon the point contact deformations at the individual balls of an angular contact ball-bearing assembly. A simulator is also developed for simulating spindle end deflections for various spindle rotational speeds. The results of the simulation show contact angle variations and peak deflections at particular spindle rotational speeds. Important research issues are also presented.Nomenclature AF final position, inner raceway groove centre - RF initial position, inner raceway groove centre - W final position of ball centre - V initial position of ball centre - D ball diameter, mm - r _o inner raceway groove radius, mm - r _i inner raceway groove radius, mm - M gyroscopic moment, N-mm - FO r _o/D - FI r _i/D - P bearing pitch diameter, mm - K _o outer race load-deflection constant, N/mm^1.5 - K _i inner race load-deflection constant, N/mm^1.5 - CF centrifugal force, N - J mass moment of inertia, N.mm² - l length of spindle, mm - E modulus of elasticity, N/mm² - I moment of inertia of spindle, mm⁴ - Y deflection of spindle alongy-direction, mm - z deflection of spindle alongz-direction, mm - M moment at spindle end, N.mm - V shear force at spindle end, N - m spindle mass, kg - material density - _o outer race contact angle - _i inner race contact angle - nominal contact angle - _i inner race deformation - _o outer race deformation - angle between ball centre of rotation and the horizontal - mis-alignment (in degrees) of shaft assembly measured in a plane perpendicular to shaft axis (x-direction) - W1 ball and raceway angular raceway velocity ratio for outer raceway control - W2 ball orbital and angular raceway velocity ratio for rotating inner raceway and outer raceway control - circumferential ball position - raceway control parameter     

The optimisation of the grinding of silicon carbide with diamond wheels using genetic algorithms

Modelling and optimisation are necessary for the control of any process to achieve improved product quality, high productivity and low cost. The grinding of silicon carbide is difficult because of its low fracture toughness, making it very sensitive to cracking. The efficient grinding of high performance ceramics involves the selection of operating parameters to maximise the MRR while maintaining the required surface finish and limiting surface damage. In the present work, experimental studies have been carried out to obtain optimum conditions for silicon carbide grinding. The effect of wheel grit size and grinding parameters such as wheel depth of cut and work feed rate on the surface roughness and damage are investigated. The significance of these parameters, on the surface roughness and the number of flaws, has been established using the analysis of variance. Mathematical models have also been developed for estimating the surface roughness and the number of flaws on the basis of experimental results. The optimisation of silicon carbide grinding has been carried out using genetic algorithms to obtain a maximum MRR with reference to surface finish and damage.Nomenclature C constant in mathematical model - C₁ constant in surface roughness model - C₂ constant in the number of flaws model - d depth of cut, m - dof degrees of freedom - f table feed rate, mm/min - M grit size (mesh) - MRR material removal rate, mm³/mm width-min - N_c number of flaws measured - R_a surface roughness measured, m - Y machining response - depth of cut exponent in mathematical model - ₁ depth of cut exponent in surface roughness model - ₂ depth of cut exponent in number of flaws model - feed rate exponent in mathematical model - ₁ feed rate exponent in surface roughness model - ₂ feed rate exponent in number of flaws model - grit size exponent in mathematical model - ₁ grit size exponent in surface roughness model - ₂ grit size exponent in number of flaws model     

Drill and clamping interface in high-performance drilling

The behaviour of a drill and a clamping unit was investigated in high-performance drilling. Some clamping units were characterised experimentally. In a series of experiments, the free-rotating drill behaviour, and the drilling events were investigated under high-performance conditions. A non-rotating measurement system, including proper procedures for signal processing, enabled the presentation of all measured values in terms and coordinates of the rotating tool. This led to a better understanding of the first-contact event, the penetration and the full drilling phases, as well as the influence of the clamping unit under different cutting conditions.Notation F impulse test exciting force [N] - Fz drilling axial force [N] - F _x F _y drilling lateral force components [N] - F _T drilling table speed (mm min^–1) - L drill overhang - T drilling torque [Nm] - X, Y, Z world coordinates [mm] - X _T,Y _T,Z _T rotating tool coordinates [mm] - _L hole location error [mm] - drill diameter [mm] - rotating angle [°] - R drill end circular movement fadius in world coordinates [mm] - X, Y drill end deflection in world coordinates [mm] - X _T, Y _T drill end deflection in world coordinates [mm] =2R     

Fast evaluation of cutting forces in milling, applying no approximate models

A new method for fast evaluation of cutting forces in milling is introduced and tested experimentally. Unlike all existing procedures, which include the use of cutting models and approximate assumptions, in this method, the elementary functions of the cutting force are obtained from measured values only.The basic force functions for the whole feed range are acquired from one experiment using a single-tooth full-diameter (slot) milling, applying a specially developed procedure. The milling experiment is conducted under low-impact conditions, enabling accurate measurement and convenient signal processing. The basic force functions are then integrated and superimposed, using known procedures, to combine the total force in any multitooth milling combination. In this work the method is explained and tested experimentally.The suggested method enables a reliable evaluation of the cutting forces, while demanding minimal experimental work, the method applies to cutters having complicated edge geometry, and to high speed milling.Nomenclature a radial depth of cut 0<a<D - feed per tooth ratio 0<1 - d axial depth of cut - D cutter diameter - a/D radial depth ratio - cutter rotation angle - cutter rotation angle [6] - F _x,y,z() instantaneous edge cutting forces in fixture coordinates - F _t,r,z() instantaneous edge cutting forces in tool coordinates - F _x,y,z ^* F_t,r,z tool cutting force components on a multitooth cutter - h instantaneous chip thickness [6] - h* equivalent edge coefficient [6] - r ₁,r ₂ tangential radial ratio coefficient [6] - K _T tangential specific cutting force [4] - K _R radial specific cutting force [4] - N number of teeth - R _r resolution reduction factor - t instantaneous chip thickness - S ₁,S feed per tooth     

Burr detection by using vision image

In this paper, a practical force model for the deburring process is first presented. It will be shown that the force model is more general than Kazerooni's model and it is suitable for both upcut and down-cut grinding. In terms of this force model, an algorithm of burr detection by using a 2D vision image is proposed. In the burr detection algorithm, the relevant data of burrs, such as frequency, cross-section area, and height are simplified so that they are functions of the burr contour only. Then, a fast tracking method of the burr contour (BCTM) is developed to obtain the contour data. Experiments show that the BCTM of this passive (i.e. without lighting) image system can be as fast as 18.2 Hz and its precision is 0.02 mm, so online burr detection and control by using the vision sensor is feasible.Nomenclature A _burr cross-section area of the burr - A _chamfer cross-section area of the chamfer - A _n proportional factor - A _work cross section area in the contact zone while deburringA _work=A _burr+A _chamfer - w cutting width - w _root thickness of the root of the burr - a depth of cut - a _root burr heighta _root=a(w _root) - C ₁ static cutting edge density - D equivalent wheel diameter - d _s wheel diameter - d _w workpiece diameterD=d _w d _s/(d _w±d _s)D=d _s andd _w for the deburring process - F _h horizontal grinding force - F _v vertical grinding force - F _n normal grinding force - F _t tangential grinding force - F _n(K) normal grinding force of the Kazerooni's model - F _t(K) tangential grinding force of the Kazerooni's model - F _o threshold thrust force - f _burr burr frequency - f _n normal grinding force per active grain - f _t tangential grinding force per active grain - f _r first resonant frequency of the robot - f _tool resonant frequency of the end-effector at the normal direction - exponential constant for describing the edge distribution = [(1 +n) + (1 –n)]/2 = (1 +n)/2 for = 0 [21] - K proportional factor of the force model of the grinding processK =A _n ^1–n / - K ₀ specific contact force per contact length - K ₁ specific chip formation force per contact length - V _s wheel speed - V _w workpiece speed - _w metal-removal parameter - K ₂ specific metal-removal parameter per wheel speedK ₂ = _w/V _s - K _c specific chip formation force per area - K _f specific friction force per area - k constant for the parabolic burr - k ₁,k ₂,k ₃,k ₄ constants for the circular burr - L contact width between the wheel and the workpieceL is equal to the chamfer's hypotenuse length, orL=w _root when there is no chamfer - l contact length - l _k contact length between the wheel and the workpiece - m exponential constant for describing the edge shape 0m1m=1 for the deburring process [21] - N _dyn number of engaged cutting edges per wheel surface - n exponential constant for describing the cutting process 0n1n=1 for the pure chip formation process andn=0 for the pure friction process [22] - average contact pressure - p exponential constant for describing the relationship between the static cutting edge and the wheel surface depth 1p2p=1 for linear case [21] - Q magnitude of the individual chip cross-section in the contact zone - r radius of the circular burr - Z _w metal-removal rate - ,, exponential constants for describing the edge distribution [21] = (p –m)/(p + 1) = 0 form = 1,p = 1 =p/(p) + 1 = 1/2 forp = 1 = (1 –n) = 1n/2 for = 1/2 - actual contact area between the wheel and the workpiece - coefficient of the sliding friction - variable of the contact angle - _k maximum contact angle - _m mean rotating angle - _t half of the tip angle of the grains - ratio of tangential chip formation force to the normal chip formation force. Usuihideji has pointed out that = /(4tan_t) [29]     

An economic design for variable sampling interval MA control charts

Most of the studies done on the economic design of control charts focus on a fixed-sampling interval (FSI); however, it has been discovered that variable-sampling-interval (VSI) control charts are substantially quicker in detecting shifts in the process than FSI control charts due to a higher frequency in the sampling rate when a sample statistic shows some indication of a process change. In this paper, an economic design for a VSI moving average (MA) control chart is proposed. The results of a numerical example adopted from an actual case indicate that the loss cost of VSI MA control charts is consistently lower than that of the FSI scheme.Design variables n Sampling size for each moving plot - h_a Subsequent sampling interval when preceding sample mean is located at sub-control region I_a, a=1,2,..., - Number of different sampling-interval lengths, 2 - k_a Threshold limit expressed in units of - k₁ Control limit expressed in units of Parameters related to assignable cause µ₀ Target mean - True-process standard deviation - Magnitude of an assignable cause expressed in units of - Occurrence rate of an assignable cause per unit timeCost and technical parameters D Average time taken to find and repair an assignable cause after detection - e Time for a sample to be taken, transmitted to laboratory, and results phoned back to process control room - M Income reduction when =₀+ - T Average cost of looking for an assignable cause when a false alarm occurs - W Average cost of looking for and repairing an assignable cause when one does exist - F_c Fixed cost per subgroup of sampling, inspecting, evaluating and plotting - V_c Variable cost per subgroup of sampling, inspecting, evaluating and plotting     

Thermographic characterisation of a laser surface engineered ceramic coating on metal

Laser-material interactions consist of complex, and generally short-lived, but intense events. Hence, many important aspects and effects of these interactions are not directly measurable, such as temperature distributions within the material. In the present study, the effect of temperature distribution on the residual stresses developed during laser surface engineering of ceramic composite coating on metal has been investigated. Infrared thermography technique has been employed as a means to measure the temperature distribution within the substrate while the laser beam is directed at the surface of the coating. Temperature distribution is generally a function of the laser input parameters, such as the laser beam power and the traverse velocity of the beam. Hence, variation in the temperature distribution and the consequent stresses developed within the composite coating due to the changing input parameters have also been investigated. The rapid processing in complement with precise control of the process based on in-situ thermographic measurements provides numerous opportunities for a high power laser as a advanced manufacturing tool.List of symbols Azimuthal angle - Tilt Angle - Q Heat energy - t Time - T Temperature - k Thermal conductivity - Density - C_p Heat capacity - Strain - d Interplanar spacing - Poissons ratio - E Youngs modulus - Stress     

An Electromagnetic Calorimeter with a Transversal Orientation of the Scintillating Fibers

The space and energy resolutions of the SPACAL electromagnetic calorimeter with a transverse orientation of the scintillating fibers have been measured. The main parameters of the calorimeter are presented. The results of experiments with an electron test beam of the DESY synchrotron at energies E = 1–6 GeV are discussed. The measured energy resolution of the calorimeter is found to be (/E) = 12.7%/E 2.0% (E is expressed in terms of GeV), and the space resolution is _{x, y} (E) = 1 mm at E = 4 GeV.     

A Modern Solution to Problems of Eddy-Current Structuroscopy

The problems arising during nondestructive quality tests of articles of nonmagnetic alloys and during sorting in the electrical conductivity are considered. The function and performance characteristics of a -26 eddy-current structruroscope are described. The structural diagram of this device and the algorithm of its operation are described. The capabilities of the -26 eddy-current structruroscope for performing nondestructive quality tests of thermal treatment of articles made of nonmagnetic materials are analyzed.     

Determining the Surface Electrostatic Potential Ψ s of a Dielectric-Bordering Semiconductor Using the Method of Ψ s "/Ψ s Diagrams

A quantitative approach to determining the integration constant _s0(V _{g 0}) in an expression relating the semiconductor surface potential _s to the voltage V _g applied to the metal–insulator–semiconductor (MIS) structure and its quasi-static capacitance–voltage characteristic C _v(V _g) (normalized to the dielectric capacitance) is described. The method is based on the analysis of experimental functions _s ^"( _s ), where _s ^" = d _s /dV _g, and the same functions calculated for an ideal MIS structure. The obtained function _s (V _g) is a rather exact and complete characteristic of electron properties of the MIS-structure phase boundary (the integrated interface state density, flat-band voltage V _FB, sign and density of the dielectric fixed charge, and variations of these parameters under the action of various factors). Using the example of a particular n-Si MIS structure, it is shown that the method of _s ^"/ _s diagrams ensures a noticeable (up to 0.93 eV) widening of the Si gap sounding region and observation (by the value of the V _FB shift) of very small ( 1 × 10⁷ cm^–2) variations in the charge density at the Si/SiO₂ phase boundary.     

The СПРУТ-VI Instrument Complex for the Mir Orbital Station

The basic elements of the instrument package program for orbital stations are presented. The characteristics of the -VI equipment developed for the Mir orbital station within the framework of this program are described. This equipment allows the simultaneous recording of the characteristics of near-earth space (electron, proton, and nucleus fluxes, magnetic fields, and low-frequency electromagnetic waves) and their effect on the elements of spaceborne equipment and systems. The mass of the equipment is 16 kg; the power consumption is no more than 20 W.     

Fracture mechanics-based model of abrasive waterjet cutting for brittle materials

Advanced engineering ceramic materials such as silicon carbides and silicon nitride have been used in many engineering applications. The abrasive waterjet is becoming the most recent cutting technique of such materials because of its inherent advantages.In the present study, two elastic-plastic erosion models are adopted to develop an abrasive waterjet model for cutting brittle materials. As a result, two cutting models based on fracture mechanics are derived and introduced. The suggested models predict the maximum depth of cut of the target material as a function of the fracture toughness and hardness as well as the process parameters.It is found that both models predict the same depth of cut within a maximum of 11%, for the practical range of process parameters used in the present study. The maximum depth of cut predicted by the suggested models are compared with published experimental results for three types of ceramics. The effect of process parameters on the maximum depth of cut for a given ceramic material is also studied and compared with experimental work. The comparison reveals that there is a good agreement between the models' predictions and experimental results, where the difference between the predicted and experimental value of the maximum depth of cut is found to be an average value of 10%.Nomenclature C abrasive efficiency factor, see equation (16) - C ₁,C ₂ c ₁/4/3, c₂/4/3 - c ₁,c ₂ erosion models constants, see equations (1) and (2) - d _a local effective jet diameter - d _j nozzle diameter - d _S infinitesimal length along the kerf - f ₁ ( _E ) function defined by equation (7) - f ₂ ( _E ) function defined by equation (8) - f ₃ ( _e ) function defined by equation (14) - g ₁ ( _E ) f ₁( _e )/f ₃ ² ( _e ) - g ₂ ( _e ) f ₂( _e /f ₃ ² ( _e ) - H Vickers hardness of the target material - h maximum depth of cut - K _c fracture toughness of target material - k kerf constant - M linear removal rate, dh/dt - m mass of a single particle - abrasive mass flow rate - water mass flow rate - P water pressure - Q total material removal rate, see equation (11) - R abrasive to water mass flow rates - r particle radius - S kerf length - u traverse speed - V material volume removal rate (erosion rate) - V idealised volume removal by an individual abrasive particle - particle impact velocity - ₀ initial abrasive particle velocity - x,y kerf coordinates - local kerf angle, Fig. 1 - _E jet exit angle at the bottom of the workpiece, Fig. 1 - particle density - _w water density On leave from: Mechanical Engineering Department, Suez Canal University, Egypt.On leave from: Mechanical Power Engineering Department, Alexandria University, Egypt.     

A Self-Initiated Periodically Pulsed Light Source on Chlorine and Krypton Chloride Molecules

An electric-discharge light source, operating in the spectral range of 170–270 nm on a system of bands of Cl₂ ( = 200 and 257 nm) and KrCl ( = 222 nm) molecules is described. The radiator is pumped by a low-pressure volume discharge in a spherical anode-flat cathode system of electrodes with an interelectrode distance of 6 cm, so that the plasma has no contact with the quartz envelope of the lamp. The working mixtures are P(Kr)/P(Cl₂) = (40–640)/(40–280) Pa. When a dc voltage U 1 kV is applied to the discharge gap, a volume discharge exists only in a periodically pulsed mode (f = 0.1–50 kHz) and represents a source of short-wave radiation with a cylindrical working surface (1 cm in diameter and 6 cm long) and a mean radiation power of 3 W.     

Desorption Kinetics of Polyether Lubricants from Surfaces

Desorption or evaporation is one of the mechanisms for loss of perfluoropolyalkylether (PFPE) lubricants from the surfaces of data storage media. One approach to minimizing PFPE loss to desorption is the use of lubricants with increasing molecular weight or increasing average chain length. In order to understand the effects of chain length on the lubricant evaporation kinetics we have studied the desorption kinetics of monolayer films of oligomeric ethers with varying chain length adsorbed on the surface of graphite. The desorption pre-exponents, v, and desorption barriers, E _des , have been measured for poly(ethylene glycol) dimethyl ethers, CH₃O(CH₂CH₂O) _m CH₃, with m=1,2,3,4,8 and 10. These are models for the PFPE known as Fomblin Z, which has a structure CF₃O(CF₂CF₂O) _x (CF₂O) _y CF₃. The results show that the desorption pre-exponents are independent of chain length and have an average value of v=10^18.7±0.3 s^–1. The E _des for the poly(ethylene glycol) dimethyl ethers vary non-linearly with chain length and can be fit with a power law expression of the form E _des =a+bN , where N is the total number of atoms in the oligomer backbone (N=3m+3) and the scaling exponent has a value of 1/2. This non-linear dependence of E _des on chain length has also been observed in recent studies of the desorption kinetics of straight chain alkanes from graphite. A desorption mechanism is described that explains the non-linearity of E _des for the poly(ethylene glycol) dimethyl ethers. The implication for the lifetime of lubricants on data storage media is that the long chain PFPE lubricants desorb more rapidly than one might expect based on simple linear scaling of the E _des of lower molecular weight PFPEs.