The effect of deoxidation of steel on machinability

Wear, 38 (1976) 1 - 16 @ Elsevier Sequoia S. A., Lauaanne - Printed in the Netherlands

THE EFFECT OF DEOXIDATION

OF STEEL ON MACHINABILITY

Z. PALMAI Iv. tit 5, 3600 Ozd - I (Hungary) (Received January 7, 1975; in final form September 2, 1975)

Summary The effect of steel deoxidation products on tool wear in machining has been studied. The mechanism of formation and the role of non-metallic protective layers on tools were investigated. The speed limit of formation of a non-metallic layer was found to be an important parameter; heat treatment can influence this. Metallurgical investigations have aided the elucidation of the deleterious effect on tool wear of aluminium deoxidation of steel.

Introduction In connection with machining technology during the last two decades several studies involving the measurement of tool life have been made. The results showed that under similar conditions tool life varied considerably [ 11. Differences found in the life of tools were attributed in most cases to the deoxidation procedure during production of the machined steel and to the characteristics of the deoxidation products remaining in the steel [ 2, 31. It was established that these deoxidation products form non-metallic layers on carbide tools and retard wear. Thus a steel-making technology could be developed to produce deoxidation products to form the non-metallic layer that retards wear [4, 51. The effect of free cutting steels of this type is shown in Fig. 1. In comparative tests, the formation of a wear crater (Fig. l(a)) on the face of the tool by turning normal steels is completely inhibited by the non-metallic layer (Fig. l(b)). A second advantage of such steels is that in the course of their machining scatter in the wear rates of the tools is significantly lower than those experienced in machining traditionally produced steels. This is shown in Fig. 2 on the basis of statistical abrasion testing. Specimens of steel charges of narrow chemical composition were turned for 5 min; the scatter of the crater depths on the tool is shown. The wide range of scatter associated with traditional steels (a) is significantly narrowed with specially produced steels (b). The machinability of steels producing non-metallic protective layers has been extensively reported [2 - 51. The cause of the wide range of scatter in Fig. 2 with traditional steels has great practical significance. Modem production methods require good machinability characteristics with low tool wear, and uniform machinability is becoming more important.

80 70

40 30 20 15 10 5 2

Fig. 1. The face of the cutting tool (a) after machining traditionally deoxidized steel, (b) after machining specially deoxidized steel. Fig. 2. Tool crater depths in machining specimens of different charges of steels of hardness ZfV = 186 f 19 kp mm-’ after turning: (a) HM steel; (b) traditional steel. Cutting data: u = 200 m min-l, T = 5 min, sa = 0.3 x 2 mm2; tool, P 20.

The present study examines the effect of the steel deoxidation products on tool wear even if no continuous non-metallic protective layer develops on the tool, as the effect of deoxidation procedure significantly influences tool life. 1. The formation of a non-metallic layer During cutting a flow zone develops on the face of the tool causing the material of the chip to slip, not at the edge of the tool but at a distance x1 from it (Fig. 3). In, the range x < x1 the speed gradient perpendicular to the surface of the tool on the face of the tool (z = 0) is a continuous function. In the range x > x1 the speed gradient at z = 0 is not continuous and the elementary particles of the chip in contact with the tool surface move with a speed u # 0. At x = x2 where the contact between the tool and the chip breaks, the speed of the flow zone at each of its points is the same as the speed up of the chip; du/dz = 0. In the range x < x1 of tke drift zone it may be assumed that du/dz = constant # 0 and the shear stress T = T(Z) = constant. On this basis, the effect

3

~

Non-metallic

layer

Fig. 3. Flow of material in chip formation. Fig. 4. The effect of an inclusion on the speed of the flow zone in the range x < x1 (Fig. 3).

of an inclusion entering the flow zone can be determined. Using the symbols of Fig. 4 assume that an inclusion of thickness AZ in the range x < xl at a distance z. < h - AZ from the surface of the tool moves together with the drifting material of the chip in the x direction. If it is supposed that at a distance z = z. + AZ measured from the face of the tool the inclusion does not influence the flow significantly then dv dv dv z,+Az)=AZ 20 + (1) dz(a” dz steel &llCl from which, in the steel layer of thickness z. moving on the face of the tool, the medium speed gradient is du du + E du du -=_ --dz steel d%” zo i d& d&i 1 Applying the assumption for the flow zone that

(2)

r = rj dv/dz

(3)

where n is the viscosity

of the material,

eqn. (2) will have the form

du du 1 -=_ +“” --- 1 1 dz steel d%” % ( %eel l)hCl Depending on the viscosity of the steel and on the inclusion flow zone, two cases can be differentiated: (a)

Vinci <

%teel

(b)

~incl>

Vsteel

(4) moving in the

In case (a), drawn with a full line in Fig. 4, dv du > -de steel dz,”

(5)

,

1000

1200

1.400

EL

Temperature PC 1

Fig. 5. The deformability of inclusions compared with that of steel.

The layer under the inclusion must suffer a deformation greater than average. If, from the point of view of deformation, limitations in the y direction are considered, the material at z < z. under the inclusion is exposed to a greater than average stress. The greater the difference between the ratio A&z0 and the shearing stress 7, or respectively the difference between the viscosities of the inclusion and the steel, the greater is the probability of the material not being able to bear the stress and of the chip splitting. The split chip in moving into the range x1 < x < 3c2rubs to the surface of the tool, the crack may widen and the inclusion can be spread as a lubricant on the surface of the tool. In case (b), shown with a broken line in Fig. 4, du du -<--(6) dz steel d%” and a reversed relation develops. The stress between the material and the inclusion is less than the average stress in the flow zone. Thus there is only a minimum probability of the chip splitting and the inclusion behaves during the process as an internal lubricant. Inclusions may get to the surface of the tool in a specific temperature interval, or in a range of cutting speed in which assumption (a) is realized. The relative plasticity of inclusions in hotly deformed steels has been characterized by Kiessling [cj] (Fig. 5). The hatched zone is where u is less than unity and meets the requirements of assumption (a). Different temperature intervals are associated with inclusions of different compositions.

5

If the roughness of the face of the tool and the effect of high stress and the complicated material structure on the viscosity are considered, the postulated mechanisms of inclusion lubrication are possible. It should be noted, however, that with inclusions of high viscosity or rigidity splitting of the bottom of the chip is of no avail as such inclusions are unable to lubricate the tool. Killed steels always contain inclusions which can become either partially or completely plastic under chip formation conditions. Thus in cutting any killed steel a non-metallic layer should develop in a continuous form or in the form of stains whenever the cutting speed exceeds a minimum boundary value which is dependent on the circumstances. The lubricating inclusion on the tool is also worn by the chip that moves on it. A continuous non-metallic layer develops if inclusion replacement occurs before it is worn away. Call the concentration of the deoxidized products which are suitable to form a protective layer c,,. The average length L of the band to be cut before another inclusion reaches the tool is (7) where pox is the density of the oxide, he is the density of the chip material, &id is the medium length of the oxide inclusions measured in the direction of cutting and To, is the oxide area on the microsection expressed as a percentage. In this way the time intervals t,,, in which successive inclusions lubricate the tool in a definite longitudinal section of the chip formation can be assumed to be

(8) The duration of the above-mentioned time intervals may vary over a wide range because of the scatter in the distribution of the inclusions. To confirm the postulates the deformation of the deoxidation products within the flow zone during chip formation was studied on longitudinal sections of chips. It was found that the plastic inclusions can be greatly deformed within the steel base material (Fig. 6(a)) but cannot reach the surface of the chip. The rigid inclusions (nine, >> nsteel) [ 51 probably get to the boundary surfaces of tool and chip from the flow layer of the chip. However, because the rigid oxide does not spread to lubricate the tool, the chip sweeps it away as indicated in Fig. 6(b). The inclusions that may be deposited on the face of the tool are less viscous than the steel but are plastic under the conditions of cutting. Such an inclusion is shown in Fig. 6(c). Only a small part of the inclusion material reaches the surface of the tool from the chip. The mean value of the time intervals GUcin which inclusions succeed each other can be estimated on the basis of an examination of the inclusions.

6

0,lmm

Fig. 6. Behaviour of inclusions of different of the chip sliding on the tool.

heat resistance

in the vicinity

of the surface

Fig. 7. The wear rate of layers in the course of turning steels which do not form a layer. Machining data, se = 0.26 X 2.5 mm2; material, C45; tool, P20. Fig._: Traces of oxide layer after turning for 1 s (250x min , sa = 0.26 X 2.5 mm ,; material, C60; tool, P20.

). Machining

data: u = 150 mm

From the investigations with a specially deoxidized steel, using eqn. (8) with &id = 7.5 X 10e3 mm, To, = 0.0167% and u = 2.5 X lo3 mm s-l, the mean value of time intervals in which inclusions succeeded each other was 7.5 x 10-S t

‘“’

=

2.5 X lo3 X 1.67 X 1W5

= 0.18 - 0.2 s

Electron

containing

QOlmm

,

,

Mn

Co

Fig. 9. An inclusion

Al

S/

mlcrogmph

rigid and plastic phases.

The wear rate of the non-metallic protective layer was also examined. Figure 7 shows the summarized measurements. It may be concluded that the time of wearing of a single inclusion lubricating the tool is of the order of seconds. From our results the thickness h of a lubricating stain on the tool formed from inclusions of average size is, for u = 150 m min-’ for example, l/3 E.tm,and according to Fig. 7 dh/dt = 10 pm min-I. Thus t

h

=-$-/2s wear dh/dt

It can be seen that in the case of specially deoxidized steels tsuc << ear and hence the non-metallic protective layer develops very quickly. By 4%.

using a tool preheated to production temperature this layer can be recognized after cutting for 1 s (Fig. 8). With traditionally manufactured steels and especially with steels doped with Al, the T,, proportion of the plastic inclusions is significantly smaller. Figure 9 shows an inclusion where only the outer shell was plastic, from which the two extensions developed. Continuous formation of a continuous non-metallic protective layer cannot be expected but from time to time areas of the layer develop. 2. The effect of oxide inclusions on the wear of the tool According to the plasticity of the inclusions of the machined steel, crater wear of tools can be characterized in the following way on the basis of Fig. 10.

Fig. 10. The density function L?ar of lubricating inclusions

(4 L

of the succeeding time t,, on the face of the tool.

<< bear: a non-metallic

continuous

and that of the wearing time

layer develops

and no crater

wear. (b) L, 2 Lear: the layer develops from time to time; the wear rate of the tool is very small. Cc) LC > twear: the micro-size stains developing from time to time can only decrease the wear rate of the face of the tool. It is usual to describe tool wear by the tool life T, which can be described as a function of the cutting speed by Taylor’s empirical formula T = C,/vn

(9)

G and n are constants.

Previously it was stated that crater wear of the carbide tool can be described by the diffusion function of tool life log T = log A + Bv-O.~

(10)

and the wear rate function dKT

log -

dt

= log K’?’ = log Al - Bv- o.2

(11)

Here B=

Q/RCln 10

A = KTb/A1

(12) (13)

9

i

-9

-6

j

m

Kol&y

0

Kmnczy

+

Kcimg

i

”

(4

.

.

Km/g

- Dedrch

0

Tdth

- Vad&

8

own

experfments

-7

-6

-5

-4 + -3

_^

(b)

Fig. 11. The relation between the Taylor formula constants CT and n and the diffusion wear function constants Q and A.

KT is the crater depth, KTb is the tool life criterion, Q the apparent activation energy of wear, R the universal gas constant, C a temperature coefficient and A, Al are quantities depending on the feature of the diffusion [ 71.

The diffusion wear function is especially suitable for examination of the effect of inclusions on wear. Considering, however, that the tool wear is usually described by Taylor’s formula, it was important to study the relation of the two tool life functions. Comparing the data obtained from the literature [8 - 131 and the present measurements, it was found that there is a close relation between C,, n and Q, A, the constants of the two functions (Fig. 11). It should also be noted that in turning carbon steels of different hardnesses the apparent activation energy Q for wear changes (Fig. 12):

70

60

50

40

30 3 >h

HV L kp/mm]

v

“bl = vbh

Fig. 12. The relation between of the machined steel.

the apparent

activation

energy

of the wear and the hardness

Fig. 13. The change of the wear rate of the tool as a function of the cutting speed: (1) machining traditionally deoxidized steels; (2) machining specially deoxidized steels.

Q = ijHV”

cal mole1 K-l

where a = 522, a = 0.86 and HV is the Vickers hardness. The effect of the inclusions on the wear rate is shown Machining traditional steel, wear curve 1, in its linear phase case (c). The greater the difference t,,,.-- twear, i.e. the area in Fig. 10, the nearer is the real wear rate to the theoretical K’Treal= A”K’T

(14) in Fig. 13. conforms to under curve A* rate: (15)

A typical example of the non-metallic layer retarding wear is shown in Fig. 14. From the raster electron micrograph, the viscous non-metallic layer can be recognized on the surface of the wear crater. Decreasing the cutting speed to the value of a dynamic balance t,,, = a boundary cutting speed us1can be specified which characterizes 44C?a*, version (b). Cutting with a slower speed than this shows that the intensity of wear significantly decreases (case (a)). On the wear curve denoted by 1 a value of the boundary speed tj,h can be designated where increased wear by high speed begins. A feature of the specially produced steels which form non-metallic layers is that these boundaries move towards an upper value (curve 2 in Fig. 13). The intensity of the wear can be basically influenced by the processes taking place in the boundary layers of the tool. This occurs in the case of a solution of WC particles of hard metal containing small quantities of Ti, Ta or C, and also on the u > ubh part of the curve (a) in Fig. 13. In these cases destruction of the tool occurs in the course of lubrication (t,“, < twear) and

1

Fig. 14. Traces of oxide layers.

following reduction of their heat resistance the partic!es are lubricated by the chips. The type of wear is similar but its intensity due to thermal insulation of the layer and the layer’s effect in decreasing the outer diffusion processes is smaller. The boundary speed hl for formation of the non-metallic layer protecting the tool defines the boundary temperature 8, of cutting at which the inclusions of the machined steel are capable of forming a layer.. Using this as a basis, and in this first instance not considering the effect of the surface pressure, the impact of the different cutting characteristics upon the L+,~ and r&h boundary speeds can be derived [ 14,151. The cutting temperature is usually given by the empirical function 8 = C’v”sYHV

(16)

where C’, LX,y and z are constants which can be defined by measurement. On converting the data of several studies into absolute temperature it was found that C’ = 412 K, x = 0.20, y = 0.12 and z = 0.08 [16]. Using these values the boundary speed is (17) The results obtained in the course of cutting traditionally produced materials are shown in Fig. 15. C,, was found to be a constant and the deviation did not exceed + 4%. 3. Applications Two practical (a) Why Al decreases

problems

are analysed.

the machinability

of steel

It is well known that Al forms rigid type inclusions in steel (Fig. 16). Their abrasive effect increases flank wear. This explanation is not satisfactory

Fig. 15. The boundary speed of the lubrication as a function of feed: a = 2 mm, y = -6”)

K =.70”, T = 5 min; tool, P20; material, traditionally deoxidized steel.

Mn-

Ka

5 -K,

Fig. 16. Microprobe analysis of a rigid inclusion unable to form a layer.

for face wear however (Fig. 6). The assumption that the lack of silicate phases rather than the presence of rigid inclusions causes crater wear of carbide tools seems to be more firmly established. In such cases the steel does not contain enough plastic inclusions to form the non-metallic protective layer. In the present studies evidence was found of non-metallic layers even in the course of cutting steel deoxidized with aluminium. Figure 17 shows

13

Fig. 17. Stains produced

on cutting

steel post-deoxidized

with Al: (a) 20x ; (b) 250x.

Imp/s

600 Electron

beam

crystal mco

500

Cl=35

)%,um 007

kV

400 300 b

700 1

200

1

Twl

Fig. 18. Microprobe analysis can be found in the crater. 405mm

Fig. 19. Pins indicating flow zone (200X ).

P20

of the tool shown

welding

in Fig. 17; localized

stains containing

45 mm

of the chip to the tool on the bottom

surface

of the

Ca

(b) Fig. 20. The effect of steel hardness: (a) on the boundary speed of lubrication; (b) in eutting a tempered specimen. Cutting data, sa = 0.42 X 2.5 mm’; material, CSO; tool, PZO.

lubricating stains on the tool. Microprobe analysis (Fig. 18) showed only Ca on the surface of the crater from the inclusions. The great variation of intensity measured at different points also indicates a stain-like arrangement. The non-covered hard metal comes into contact with the material of the chip and localized welding occurs. The shearing of these welds accelerates tool wear as can be seen in Fig. 19, which shows the effect of knocking the tool downwards out of its place ; brbken welds are indicated. The deleterious effect of Al has been reduced by special deoxidation with Ca [17,181. In this way the rigid type of inclusions ensuring the fine granular structure of tbe steel can be combined with plastic phases which improve machinability, as shown in Fig. 9. (The plastic phase shown in this figure is silicate and not sulphide [ 181,) (b) How the heat treatment of steel influences the boundary speed of formation of the ~~~t~cti~e layer Boundary speed, marked u,, in Fig. 13, is irnpo~~t in both tradition~ly and specially produced steels because cutting these with a speed q, > u causes considerable tool wear. Using a tempered specially deoxidized steel bar of C 60 quality with a diameter of 110 mm, the boundary speed v,, was determined as a function of hardness (Fig. 20). From the results the boundary temperature value ch~ac~~stic of the fo~ation of the non-me~lic protective layer was computed by using eqn. (16). It was found that Bb = 1596 K * 3%. The boundary speed as a function of the hardness of the steel is shown in Fig. 21. It appeared useful to study the relation between the boundary speed and the diffusion theory of tool wear.

15

Fig. 21. The speed limit of lubrication as a function of hardness.

270

250

230

Fig. 22. Analysis on the basis of eqn. (18) of the factors influencing the speed limit of lubrication.

Substituting C” = C’sy , gives

eqns. (14) and (16) into the wear function

log Ki;, = log Al -

Q RC”

(ll),

where

HV0.78 - u0.20

Here the critical speed of wear K$, (Fig. 13) is the limit of the severely accelerated wear and it can be considered constant.

(18)

16

If within the above-studied phenomena Al = AI and other factors can be excluded, it may be assumed on the basis of eqn. (18) that there must be a function HV =

f

(HV”.78/v”20)

(19)

This assumption is confirmed by Fig. 22. The measured points are very close considering the wide deviation usually found in cutting experiments. A more comprehensive analysis could provide useful information for this development of hard metals. References 1 R. Schaumann, Streuwertuntersuchungen der Zerspanbarkeit von Stahlwerkstoffen, Maschinenmarkt, 47/48 (1956) 37 - 52. 2 H. Opitz, M. Gappisch, W. Kiinig, R. Pape and A. Wither, Einfluss oxydischer Einschliisse auf die Bearbeitbarkeit von Stahl Ck 45 mit Hartmetalldrehwerkzeugen, Arch. Eisenhiittenwes., 12 (1962) 841 - 851. 3 J. Pietikiiinen and M. Tohka, Effect of various deoxidation treatments and tool conditions on the machinability of slightly hypoeutectoid carbon steels, Jernkontorets Ann., 154 (5) (1970) 215 - 227. 4 A. Wither and R. Pape, Metallurgische Voraussetzungen fiir die Bildung oxydischer Bellge auf Hartmetallwerkzeugen bei der Zerspannung von Stahl, Stahl Eisen, 21 (1967) 1262 - 1269. 5 Z. PLlmai, Specially deoxidised free cutting steels, Met. Mater., June (1974) 326 - 330. 6 R. Kiessling, The influence of non-metallic inclusions on the properties of steel, J. Met., Oct. (1969) 48 - 54. 7 Z. Pilmai, A new physically defined function to describe the wear of cutting tools, Wear, 27 (1974) 251- 258. 8 W. Kanig, Beitrag zur Ermittlung der Ursachen fiir ein Zerspannung von Werkstoffen gleicher Normbezeichnung mit Hartmetalldrehwerkzeugen, Dissertation, Technische Hochschule, Aachen, 1962. 9 I. Kal&szi, A fGmforglsol& kop&jelendgeinek vizsghlata, Dissertation, Budapest, 1963. 10 M. Kazinczy, Tool life criteria of single point tools when cutting NC machine tools, Int. Conf. on Production Research, Birmingham, 1970. 11 T. Toth and D. VadLz, AZ optimilis Cjraelez&i szPm meghatlrozasa kisbrletianalitikai mbdszerrel kriteres szerszPmkopb esethben, Miisz. KSzlemBnyek, 41 (3 - 4) (1969) 207 - 224. 12 W. Kijnig and N. Diederich, Cutting fluids improve tool life of carbide tools by chemical reactions, Ann. C.I.R.P., XVII (1969) 17 - 28. 13 M. Szafarczyk, R. Kolendowicz and A. Winiarski, Forgicsolisi jellemziik az adaptiv szabilyozk szemszagebS1, GipgyLrt&techno&ia, XII (10) (1972) 441 - 445. 14 Z. Pfilmai, M6dszerek a jS1 forg&solhat6 szerkezeti acelok vizsgilatghoz, GBpgyirtLtechnolbgia, XII (1) (1972) 18 - 22. 15 W. Kiinig and N. Diederich, Einfluss nichtmetallischer Einschliisse auf die Zerspanbarkeit des Stahles Ck 45, Arch. Eisenhiittenwes., 41 (3) (1970) 267 - 277. 16 Z. Pglmai, AZ a&l dezoxidicibjhnak hat&a a megmunk&lhat&&gra, Dissertation, Budapest, 1973. 17 E. Miyoshi, N. Kakimi, T. Kato, M. Takoto and E. Tamura, On medium carbon-free cutting steel for machine structural use, Sumitomo Search, 3 (1970) 110 - 129. 18 V. A. Tipnis, R. A. Joseph and J. H. Boubrava, Kalziumdesoxydierte Stiihle mit verbesserter Bearbeitbarkeit fir Kraftfahrzeuggetriebe, Werkstatt Betr., 106 (12) (1973) 1017 - 1024.