Surface integrity-of-ground-components
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1) The document analyzes the surface integrity of En31 steel after grinding with different abrasives. Temperature measurements showed that tensile residual stresses formed before visible damage like oxidation or phase transformations. 2) High efficiency deep grinding was shown to reduce thermal damage potential under optimal conditions by using an inclined heat source with circular contact. 3) The choice of grinding abrasive was found to be important for controlling thermal effects on surface integrity. Controlling grinding temperature is key to preventing different types of damage.
![martensite111 ] q
T&w , , , , . ,
0 100 200 300 400
depth (vm)
Figure 3. Re-hardening burn - microhardness
Figure 3 shows microhardness readings taken below the surface of the sample shown in
figure 2. The readings confirm the presence of UTM in the surface layer. The
microhardness of the sub-surface layer has reduced to below that of the original
tempered martensite, indicating a layer of temper burn or over-tempered martensite.
OVERTEMPERING I TEMPER BURN
Tempering gives time for atoms of carbon to begin to diffuse back to their original
structure, giving rise to a relaxation of the distorted martensite lattice. If the temperature
of a previously hardened workpiece is raised above the original tempering temperature,
but below its transformation temperature, then overtempering or “temper burn” will
occur. The surface layer of the workpiece will be softened, resulting in a drop in wear
resistance of the component. Figure 4 is an example of an overtempered En31 bearing
steel. A dark layer beneath the surface, of thickness approximately lOOurn, indicates
the presence of over-tempering.
Figure 4. Temper burn](https://image.slidesharecdn.com/surface-integrity-of-ground-components-180110015505/85/Surface-integrity-of-ground-components-5-320.jpg)
![Figure 5 shows microhardness readings taken across the area shown in figure 4,
demonstrating the drop in hardness with over-tempering.
H,
I
RR
kg/mm3 Q
00
0 0
700
R
600 ]
0
t I
100 200
depth hm)
Figure 5. Temper burn - microhardness
RESIDUAL STRESSES
Grinding can be a major cause of surface residual stress in engineering components
and can therefore critically influence service life. Depending on their sign and
magnitude, surface residual stresses can either increase or decrease component
lifetimes.
In terms of the three factors influencing surface residual stresses due to grinding it was
found: -
(1) There is a direct relationship between tensile residual stress and temperature,
showing that it is thermal expansion/contraction, which is responsible for generating
tensile residual stresses in grinding.
(2) Compressive stresses, caused by mechanical effects, dominate at low
temperatures but are eliminated when thermal effects reach a critical level.
(3) Phase transformations, although capable of changing the residual stress-state,
are not relevant to the onset of tensile residual stress.
The onset of tensile residual stress is caused by exceeding a critical transition
temperature, as shown in figure 6. It was found that the transition temperature is
dependent on; the type of material being ground; the heat treatment history; the value of
yield stress and its relationship with temperature. The effects of the process parameters
such as workspeed were also shown to vary the onset temperature for tensile residual
stress. Using a critical transition temperature reduces the problem of controlling residual
stress into the easier problem of controlling grinding temperature.](https://image.slidesharecdn.com/surface-integrity-of-ground-components-180110015505/85/Surface-integrity-of-ground-components-6-320.jpg)
Surface integrity-of-ground-components
- 1. MROI -236 QL W n < n Z Society of Manufacturing Engineers 2001 Controlling the Surface Integrity of Ground Components author(s) DAVID F. McCORMACK Senior Project Engineer W. BRIAN ROWE Director of AMTTREL TAN JIN Researcher Liverpool John Moores University Liverpool, Merseyside, UK abstract The surface integrity of En31 was assessed after grinding with CBN and alumina abrasive and different types of materials damage were described. Temperature measurements were carried out for different grinding processes, these temperatures were compared with different types of material damage including burn and the formation of tensile residual stresses. It was found that the onset of tensile residual stress occurred independently of, and before, any visible damage such as oxidation, burn or phase transformations. High efficiency deep grinding (HEDG) was analyzed, using an inclined heat-source with a circular arc of contact, and was shown to give reduced potential for thermal damage under optimum conditions. The choice of grinding abrasive was shown to be an important factor for control of thermal effects on surface integrity. conference 4th INTERNATIONAL MACHINING & GRINDING May 7-10, 2001 Troy, Michigan terms Grinding Surface Integrity Residual Stress HEDG Sponsored by the Composites Manufacturing Association of the Society of Manufacturing Engineers One SME Drive P.O. Box 930 Dearborn, MI 48121 Phone (313) 271-1500 www.sme.org
- 2. D F McCormack, W B Rowe and T Jin AMTTREL, School of Engineering, Liverpool John Moores University, UK d.f.mccormack@livim.ac.uk ABSTRACT The surface integrity of En31 is assessed after grinding with CBN and alumina abrasive and different types of materials damage are described. Temperature measurements were carried out for different grinding processes, these temperatures were compared with different types of material damage including burn and the formation of tensile residual stresses. It was found that the onset of tensile residual stress occurred independently of, and before, any visible damage such as oxidation, burn or phase transformations. High efficiency deep grinding (HEDG) is analyzed, using an inclined heat-source with a circular arc of contact, and is shown to give reduced potential for thermal damage under optimum conditions. The choice of grinding abrasive is shown to be an important factor for control of thermal effects on surface integrity. KEYWORDS: Grinding, Surface Integrity, Residual Stress, HEDG INTRODUCTION Elevated temperatures in grinding can critically influence the surface integrity and life of a component. Thermal damage of a component includes burn and the formation of tensile residual stresses. Burn is a general term used to indicate that the surface of a ground component has been damaged by surface heating. The term can be used to describe several types of damage caused by high temperatures including, the formation of surface oxidation and sub-surface metallurgical changes. The nature of the material changes associated with burn depends upon the material properties and the temperature-time history to which the material has been exposed. The influence of grinding heat source and choice of grinding abrasive are important factors in maintaining surface integrity of a ground component. SURFACE OXIDATION Oxidation of the material surface occurs with exposure to high temperatures during grinding and is seen as a discoloration of the workpiece surface. The discoloration is analogous to the temper colours produced during conventional heat treatment, however, the oxidation colours produced during grinding do not occur at the same temperatures as for conventional heat treatment. The oxidation of the material surface is an Arrhenius rate controlled process. The thickness and colour of the oxide layer is dependent on time and temperature. The temperatures at which oxidation occurs during grinding, at
- 3. normal workspeeds, are high compared with conventional heat treatment. This is because the duration of the thermal pulses are very short in grinding, typically less than 20ms. During conventional heat treatment, oxidation arises from lower temperatures applied for longer periods of time. It has been found that the onset of light straw colours for a range of ground ferrous materials occurs mainly in a range between 450 and 500°C [Rowe, 19951. Severely burned ferrous workpieces are easily recognised and may be brown or even blue/black on the surface. An example of a severely burned workpiece is shown in figure 1. Dark bands are clearly visible on the surface of the workpiece. The banding is not only evidence of burn but of vibration, also associated with the onset of burn. Figure I. Surface oxidation Slight oxidation of the surface of a non-critical component may not in itself be considered detrimental to the reliable operation of that component. The oxidation can be polished off and the surface appearance restored. Oxidation of the surface of a critical component is unacceptable, due to the associated risk of sub-surface metallurgical damage and tensile residual stress formation. RE-HARDENING BURN During the grinding process, the action of the heat source moving across the workpiece surface causes rapid heating and cooling of the surface layer to take place. Re- hardening burn is the term given to the damage which occurs when the surface layer of a previously hardened and tempered ferrous component is ground so abusively that the temperature of the surface layer rises to a point necessary for the formation of austenite. When the grinding heat source has passed, the surface layer is cooled at a rate which is greater than the critical cooling rate for the material. Bulk cooling from the workpiece is rapid and causes re-hardening. Re-hardening may lead to the formation of untempered martensite (UTM), and may also give rise to the presence of small amounts of retained austenite, particularly in high alloy steels.
- 4. Hardening involves the non-equilibrium transformation of austenite. Face-centred cubic (fee) austenite will be formed within the surface layer of a workpiece, which has been subject to grinding surface temperatures in excess of the austenitizing temperature for that material. As a result of rapid cooling there is insufficient time for the fee, carbon- rich, austenite to transform back to body-centred cubic (bee) ferrite and cementite. As the temperature drops below approximately 3OO”C, a distorted lattice structure is produced known as martensite. Martensite forms by a sudden shear process in the austenite lattice, not normally accompanied by atomic diffusion, giving rise to a characteristic lenticular microstructure [Honeycombe, 19811. In grinding, this transformation takes place in less than lus because the movement of carbon atoms does not rely solely on diffusion, as in conventional heat treatment, but also on transport within the plastically deformed surface material [Shaw, 19941. Martensite is a body- centred tetragonal (bet) structure which is of increased volume compared with bee ferrite and cementite, causing an increase in lattice distortion. The increase in lattice volume with transformation to martensite causes the formation of tensile residual stresses and hardening within the surface and sub-surface layer, increasing the likelihood of crack formation. The formation of UTM may lead to premature failure of a component due to fatigue. An example of a re-hardened surface can be seen in figure 2. The material is En31 bearing steel. The material was originally hardened by oil quenching from 84O”C, followed by a sub-zero cooling process to -73°C for 2 hours, in order to transform any remaining austenite to martensite. The material was then tempered at 150°C for 2 hours to give a hardness of 850 HV. Three distinct zones can be seen in the microstructure, (i) a UTM layer at the surface, of thickness approximately 20um, (ii) an over-tempered dark layer beneath the surface, of thickness approximately 300um, (iii) a lighter coloured sub- structure, comprised of the original tempered martensite. Figure 2.Figure 2. Re-hardening burnRe-Gydening burn *
- 5. martensite111 ] q T&w , , , , . , 0 100 200 300 400 depth (vm) Figure 3. Re-hardening burn - microhardness Figure 3 shows microhardness readings taken below the surface of the sample shown in figure 2. The readings confirm the presence of UTM in the surface layer. The microhardness of the sub-surface layer has reduced to below that of the original tempered martensite, indicating a layer of temper burn or over-tempered martensite. OVERTEMPERING I TEMPER BURN Tempering gives time for atoms of carbon to begin to diffuse back to their original structure, giving rise to a relaxation of the distorted martensite lattice. If the temperature of a previously hardened workpiece is raised above the original tempering temperature, but below its transformation temperature, then overtempering or “temper burn” will occur. The surface layer of the workpiece will be softened, resulting in a drop in wear resistance of the component. Figure 4 is an example of an overtempered En31 bearing steel. A dark layer beneath the surface, of thickness approximately lOOurn, indicates the presence of over-tempering. Figure 4. Temper burn
- 6. Figure 5 shows microhardness readings taken across the area shown in figure 4, demonstrating the drop in hardness with over-tempering. H, I RR kg/mm3 Q 00 0 0 700 R 600 ] 0 t I 100 200 depth hm) Figure 5. Temper burn - microhardness RESIDUAL STRESSES Grinding can be a major cause of surface residual stress in engineering components and can therefore critically influence service life. Depending on their sign and magnitude, surface residual stresses can either increase or decrease component lifetimes. In terms of the three factors influencing surface residual stresses due to grinding it was found: - (1) There is a direct relationship between tensile residual stress and temperature, showing that it is thermal expansion/contraction, which is responsible for generating tensile residual stresses in grinding. (2) Compressive stresses, caused by mechanical effects, dominate at low temperatures but are eliminated when thermal effects reach a critical level. (3) Phase transformations, although capable of changing the residual stress-state, are not relevant to the onset of tensile residual stress. The onset of tensile residual stress is caused by exceeding a critical transition temperature, as shown in figure 6. It was found that the transition temperature is dependent on; the type of material being ground; the heat treatment history; the value of yield stress and its relationship with temperature. The effects of the process parameters such as workspeed were also shown to vary the onset temperature for tensile residual stress. Using a critical transition temperature reduces the problem of controlling residual stress into the easier problem of controlling grinding temperature.
- 7. 1000 T s 800 k ;;; 600 ! z I g 400 ’ T zJ 200 1 E u) 2 0 I- -200 y 50 AA A / AI I). 100;750l2&P 0 0 AA : Ad 0 250 300 350 -400 1 Temperature rise (“C) Figure 6. Experimental results showing residual stress versus temperature rise in En31 MEASUREMENT OF GRINDING TEMPERATURES Temperatures were measured experimentally using embedded ‘grindable’ constantan- iron thermocouples to record the temperature as the wheel passed over the junction. A single constantan pole was used, insulated by two adjacent layers of mica, located inside a pre-machined slot. The average junction width was <25pm for both the foil and wire thermocouples. Typical thermocouple assemblies are shown in figures 7 and 8. Figure 7. Foil thermocouple assembly
- 8. Figure 8. Wire thermocouple assembly.Figure 8. Wire thermocouple assembly. ANALYSIS AND EVALUATION OF RESIDUAL STRESS PROFILES Near-surface residual stress profiles were analysed using X-ray diffraction (XRD) and neutron diffraction. The XRD depth results were validated using neutron diffraction, requiring development of the neutron diffraction method. The XRD measurements were carried out by The Open University, UK, Department of Materials Engineering. The Neutron measurements were performed at the ISIS facility of the Rutherford Appleton Laboratory, UK using the ENGIN instrument. For the XRD measurements a Brucker-Siemens 05005 diffractometer was used. Residual stresses were obtained by analysing the displacement of the diffracted peaks, and qualitative information about the plastic deformation on the surface was obtained by analysing the change of the integral peak widths (IPW) [Maeder, 19811. Several methods were investigated in order to determine the peak location accurately for different peak shapes. Traditionally for simple materials such as En9, where the diffracted peak is well defined, a standard procedure for XRD peak analysis is used. The peak profile above 85% of the maximum intensity is fitted with a parabola, and the peak centre 28 taken as the maximum of this parabola, as shown in figure 9. Analysis of En31 was more complicated, owing to the much broader peaks obtained. The peak broadening is a result of the lattice strain induced by the high martensite and carbide content of these materials. For En31 the sliding gravity method was used to obtain peak centre location [Convert, 1984; Pfeiffer, 19941, where the position of the peak centroid was evaluated using progressively greater proportions of the peak data, as shown in figure IO. Measurements at each point analysed were taken by tilting along two orthogonal axes perpendicular to the beam, allowing the full stress tensor (assuming os=O) to be determined. The reproducibility of the measurements was found to be extremely good, making it possible to successfully analyse the data from this microstructurally complex material. For the in-depth stress measurements, surface removal by local electrochemical polishing was performed. The surface removal was performed on an area of 0.5cm2 in steps between IO and 50ym, making any stress relaxation negligible (Moore, 1958; Castex, 1984).
- 9. Neutron diffraction surface stress measurements were used to validate, non- destructively, the XRD stress profiles. A novel surface scanning geometry was used, which proved capable of achieving near surface measurement (400pm) pang, 1998; McCormack, 20001. Beyond an initial surface measurement, where the average stress over 5pm is measured, XRD depth measurements are destructive, as surface removal by local electrochemical polishing is required. Figure IO. En31 X-ray peak It was found that the stress in the grinding direction was always more tensile or less compressive than in the perpendicular direction. From the depth analyses it was observed that the maximum stress for the majority of the samples was present on the top surface. The residual stress affected layer varied from IOum to 300um depending on the process parameters and the abrasive type used. The affected layer depth was determined by profiles of both stress and IPW. It was found that the IPW were proportional to the density of plastic deformation in the material. Two different types of curve shapes relating to the surface stress-state were obtained from the En31 samples studied; tensile profile with maximum stress on the surface and compressive profile with maximum stress on the surface. Figure 11 shows an example of the profiles measured on En31, using both XRD and neutron diffraction. There is excellent agreement between the XRD and neutron measurements, taking into account the fact that the neutron results are the average values within the penetration depth. With the neutron technique it is almost impossible to measure the top surface point, however, the stress over the first 5pm can be obtained non-destructively using XRD.
- 10. t 100 200 Depth (pm{ -250 1 0 Figure 11. Comparison of XRD and neutron measurements on En31 CRITICAL DAMAGE TEMPERATURE Prevention of material damage during grinding requires the maximum temperature to be kept below the critical damage temperature for the material. The critical damage temperature must be specified for the particular type of damage, which it is desired to avoid. When grinding En31: - 0) In order to prevent the formation of tensile residual stresses the maximum surface temperature should be kept below 250°C figure 6. The transition temperature is dependent on; the type of material being ground; the heat treatment history; the value of yield stress and its relationship with temperature. The effects of process parameters, such as workspeed, can also vary the onset temperature for tensile residual stress. (ii) The critical damage temperature for the onset of temper burn is approximately 450°C [Rowe, 19951. When temper burn occurs, tensile residual stresses are already in evidence. (iii) The temperature at which phase transformation takes place under equilibrium conditions is known to be 723°C for plain carbon steels. This temperature may be further lowered by the addition of alloying elements. There is a danger of re-hardening burn occurring where measured grinding temperatures approach 650°C. Severe workpiece damage may be encountered due to higher than expected grinding temperatures resulting from the process becoming unstable. Table 1 Material En31 Critical Damage Temperatures Residual Stress Temper Burn Transition 250°C 450°C Re-hardening Burn 650-723°C
- 11. 10 EFFECT OF HEAT SOURCE ON TEMPERATURE Whilst increasing material removal rates give rise to increasing grinding temperatures in shallow-cut and creep grinding, reduced grinding temperatures can be achieved with increased material removal rates using HEDG (High Efficiency Deep Grinding) technology. Whilst HEDG can yield lower temperatures under optimum conditions it can also give disastrous results if these conditions are not optimised. Current research is limited and further investigation of the process is required to optimise reduced values of specific energy with increasing workspeed, increasing depths of cut and effective fluid delivery. HEDG uses a combination of deep grinding at high workspeeds and very high removal rates to give values of specific energy which, are lower than those obtained in shallow- cut and creep grinding [Rowe, 20011. Experiments were carried-out to demonstrate the high removal rates achievable and to measure the resulting contact temperatures. Experiments were conducted on an Abwood 6 kW surface grinding machine using an alumina wheel. The process conditions are detailed in Table 2. The workpieces used were made from two 0.6 mm steel blades, sandwiching a constantan wire thermocouple, as in figure 8. The samples were 38-55 mm long and 1.2 mm wide. Power was monitored during up-grinding along the length of the samples. The samples were offset at an angle of I” to the grinding direction, to give a more reliable thermocouple junction. Increased wheel wear, due to the high removal rates, was offset by the reduced specific energy and by increasing wheelspeed. A jump in temperature was observed in the transition region between boiling and burn- out of the grinding fluid, figure 12. Measured contact temperatures are compared with theoretical temperatures in figure 13. The theoretical temperatures previously published work [Rowe and Jin, 20011, using a model based heat source with a circular arc of surface. Values of specific energy are removal rate in figure 14. Table 2 HEDG Test Conditions are based on on an inclined shown against
- 12. 11 + vf =0.32 m/s 1000 j 0 Burn out 00 0 800 1 0 600 0 400 i Up to boiling Up to boiling 200 d w n o vf =0.3 m/s -Calculated mean ” rQ;, ( mm% ) 220 240 260 280 300 320 Figure 12. Measured maximum contact temperatures in near-transitional conditions e, (J/mm3) 24 Tmax (Deg.C) A vf =0.2 m/s 1500 i A A 13001 - L i A 4 . -Calculated mean 1100 i +A 63 0 vf =0.25 m/s 900 I 1700 / - / . Calculated V, =0.25 m/s 500 I 60 100 140 180 *i. Q,’ (mm%.) Figure 13. Measured maximum contact temperatures in burn-out conditions I / 70 120 170 220 270 Figure 14. Specific energy versus removal rate in HEDG EFFECT OF ABRASIVE TYPE ON TEMPERATURE The best method of achieving cool grinding is to reduce the grinding energy entering the workpiece surface, by ensuring that the partition ratio, &,, is kept to a minimum. The rise of the surface temperature of a workpiece during grinding depends upon how much of the total energy generated during grinding enters the workpiece. The thermal characteristics of the grinding wheel play an important role in energy partitioning. Due to their greater grain conductivity, CBN abrasives carry more energy away from the grinding zone than do alumina abrasives, giving a lower R, and leading to cooler grinding. Grinding with CBN leads to cooler grinding, giving decreased likelihood of either workpiece burn or the formation of thermally induced tensile residual stresses in a component surface. The effect of grinding with CBN can be seen in figure 15 for the cylindrical grinding of En31. Use of CBN abrasive has led to high magnitude compressive residual stresses being generated in the workpiece at temperatures below the critical transition
- 13. 12 temperature for En31, for removal rates between 14-57mm*/s. This is in contrast to the results obtained with alumina abrasive where extremely high temperatures and, thus, tensile residual stresses have been generated, for removal rates of only 19 mm*/s. Workpiece: En31I-Rc 54-68 Wheel: CBN (B91) 1000 T 800 t Alumina Alumina (73A601J8V) Q'w= 19mm'/s : Machine: Jonesand Shipman Series 10 CBN Q'w= 14-57 mrr?/s -800 - Temperature rise (“C) Figure 15. Experimental results showing the effect of abrasive-type on temperature and residual stress during cylindrical grinding of En31 CONCLUSIONS Temperatures were measured during grinding and related to the onset of different types of damage, including burn and the formation of tensile residual stresses, in the form of critical damage temperatures. The combination of XRD surface stress measurement and neutron diffraction sub- surface measurement has produced a technique for non-destructively mapping the residual stress profile within a ground surface. This offers significant advantages, particularly for the evaluation of the residual stress-state of expensive ‘in-service’ components within the aerospace industry. HEDG has the potential to reduce the transmission of heat to a ground finished surface due to the combined effects of large inclination angle and high workspeeds. It was shown that very efficient grinding can be achieved using HEDG under optimum conditions, with values of specific energy lower than those achieved in shallow-cut or creep grinding. The transition from boiling to burn-out of the coolant has a strong influence on contact temperatures. Grinding with CBN abrasives leads to cooler grinding and a decreased likelihood of imparting a tensile residual stress-state to the ground component surface.
- 14. 13 ACKNOWLEDGEMENTS Acknowledgements are due to Corus, EPSRC, Jones and Shipman, Timken, TRW and Unicorn Van Moppes for funding of the work on residual stresses. Acknowledgements are also due to Dr Lyndon Edwards, Dr Michael Fitzpatrick and Dr Ahmed Bouzina of The Open University, UK, Department of Materials Engineering, for their contribution to the work involved with the analysis and evaluation of the residual stress profiles. REFERENCES Castex L, Stress Redistribution after Surface Removal, French Groupement for Residual Stress Analysis Aix-En Provence 1984. Convert F, Localisation of Diffraction Peaks Using Centroid Method, French Groupement for Residual Stress Analysis Aix-En Provence, 1984. Honeycombe R.W.K, ‘Steels-Microstructure and properties’, Edward Arnold, 1981. McCormack D.F, Rowe W.B, Chen X, Bouzina A, Fitzpatrick M.E and Edwards L, Characterising the Onset of Tensile Residual Stresses in Ground Components, Proc. Sixth International Conference on Residual Stresses (ICRS-6), Oxford University, 2000. Maeder G, Castex L, LeBrun J.L and Sprauel M, Residual Stress Determination by X ray, Pub. SC. & Tech ENSAM, 22, 1981. Moore M.G and Evans W.P, Mathematical Correction for Stress in Removed Layers in X-ray Diffraction Residual Stress Analysis, SAE Transactions, Vol. 66 1958, p. 340-345. Pfeiffer W, The Role of the Peak Location Method in X-ray Stress Measurement, Proc. of the Fourth Int. Conf. on Residual Stresses, 1994, SEM, Bethel, CT, USA, 148-155. Rowe W.B, Black S.C.E, Mills B, Qi HS, and Morgan M.N, Experimental Investigation of Heat Transfer in Grinding, Annals of the CIRP, 44, 1, (1995), pp.329-332. Rowe W.B, Thermal Analysis of High Efficiency Deep Grinding, Int. J. Machine Tools and Manufacture, 41, 1, 1-l 9, 2001. Rowe W.B and Jin T, Temperatures in High Efficiency Deep Grinding, Annals of CIRP, accepted for publication, 2001. Shaw M.C, Heat-Affected Zones in Grinding Steel, Annals of the CIRP, 43, 1, (1994), pp.279-282. Wang D, Harris LB, Withers P.J. and Edwards L., Sub-surface Strain Measurement by Means of Neutron Diffraction, Proc. Fourth European Conference on Residual Stresses, Ed. Denis, S et al, Societe Francaise de Metallurgic et de Materiaux, 1, pp.69-78. 1998.