1/17
DENKI-SEIKO (Electric Furnace Steel), vol. 71 (2000), No. 2, p. 119 - 129
Paper

Effect of Si Content on the Machinability of Hot Working Die Steels
Toshimitsu Fujii and Yukinori Matsuda

1. Preface
Hot working die steels containing 5% Cr such as JIS SKD61 are usually used with 42 to 50 HRC hardness
after hardened and tempered. In recent years high performance steels featuring low Si and high Mo content
have been developed because of findings of those feature’s effectiveness to higher toughness and heat checkresistance1). These steels have been commercialized in particular for die-cast molds with design faces2), 3).
However the wider they were used, the more problem of less machinability than conventional steel like
SKD61 has been reported. Therefore countermeasures are desired. Studies on the machinability against Si
content for annealed 5% Cr steels (85 – 96 HRB) 4)or hardened and tempered 0.3C-4Cr-1.5Mo-0.6V-Ni
steels (43 HRC)5) have been investigated. As the result of those studies, machinability decrease along withdecrease of Si content have been clarified, although the mechanism behind the phenomenon has been
unexplained well enough.
Moreover recently “Direct Milling” which enables elimination of rough machining of annealed materials
by milling directly hardened and tempered steels have been widely applied for the purpose of cost or leadtime reduction in hot working die steels.6), 7) This “Direct Milling” have already been applied to 50 HRC
material area6). Therefore machinability of hardened and tempered steels rather than annealed ones andhigher hardness steels among hardened and tempered ones tend to be more important. However as
previously mentioned available reports on Si content and machinability for hot working die steels are limited
to relatively low hardness ones, less than 43 HRC.4), 5) There are no reports on the effect of Si content on
the machinability of hard steel near upper limit of usage area.
Therefore in this study using conventional 5%Cr-3%Mo steel as high performance steel, the effect of Si
content on the machinability of hot working die steels have been investigated as well as cutting mechanism
change by Si addition.

2. Test conditions
2.1 Specimen
Table 1 shows chemical compositions of specimens. Five steels containing Si from 0.05% to 1.5% havebeen prepared while main components have remained as Fe-0.35%C-0.6%Mn-5%Cr-3%Mo-0.85%V.
They were melted by Vacuum Induction Heated furnace and casted to 150 kg of ingots. Ingots were heated
to 1503 K, forged to 68 mm squared billets and annealed.

© 2020 Daido Steel Co., Ltd.

No. SD2005 (2020.08)

2/17
Table 1. Chemical compositions of test steels
steel

C

Si

Mn

Cr

Mo

V

A

0.33

0.05

0.60

5.53

3.01

0.83

B

0.33

0.15

0.60

5.53

3.00

0.84

C

0.33

0.29

0.60

5.55

3.01

0.84

D

0.34

0.99

0.60

5.55

3.01

0.84

E

0.34

1.52

0.60

5.54

3.01

0.84

Specimens for microscopic examination, end-mill machining, turning and high temperature tensile tests
were taken from annealed billets, machined roughly, hardened and tempered and finished before providingthem for each testing. All specimens were heated at 1303K for 1.8 ks, quenched in 6 bar of gas before
tempered for 3.6 ks twice at 878K for steel A and 883 K for other four steels in order to adjust hardness as
49 ± 1HRC. Si is well-known additive to increase high temperature softening resistance of steel8). Thisphenomenon was observed in these materials as well since higher Si steel had more secondary hardening
amount after tempering. However difference of hardness after tempering according to Si content aroundtested temperature, 880 K is rather small, and the difference of tempering temperature to acquire the same
hardness is as small as 5 K.
2.2 Experimental procedure(1) Microstructure observation
Microstructure were observed for sections parallel to forging directions. Surfaces of specimens were
polished and etched with Nital before optical microscopic observation were taken place. Scanning Electron
Microscope (SEM) observation with as-polished specimens were also taken place to investigate the state ofcarbides precipitation. SEM observation was also carried out for as-quenched samples.
(2) Evaluation of machinability by end-milling
Generally machinability is evaluated by 1) chip processing capacity, 2) cutting resistance, 3) tool life, and4) machining accuracy9). In this study machinability of end-milling cutting which is widely used for moldmachining was evaluated by tool life that is usually problematic in actual process. Cutting condition is shownin Table 2. Side-face cutting with 32 mm diameter throw-away end mills were taken place using #40 taper
spindle horizontal machining center. Tool life was defined as times when maximum flank face abrasions
become 300μm by measuring abrasion amount of carbide tool every certain cutting length. Specimen
dimension was 65 mm squared by 200 mm length. Measurement of tool abrasion amount was taken placeat cutting length of 2n X 100 mm. Damage state of same carbide tools under every certain cutting distancewere observed by SEM to understand damaging process. These observations were taken place for different
tools from tool life measurement tools to eliminate effect of treatment of SEM observation on tool life.

© 2020 Daido Steel Co., Ltd.

No. SD2005 (2020.08)

3/17
Table 2. Cutting conditions for tool life test in end-milling
Diameter of end-millΦ32

Number of tooth1

Tool grade

M20(Carbide), Non coating

Tool geometry

AEMW19T308ER

Axial depth of cut

4 mm

Radial depth of cut

1 mm

Cutting speed

35.8 m/min

Feed

0.15 mm/tooth

Cutting method

Down cutCutting fluid

Dry

Criterion for tool life

Flank wear = 300 μm(3) Measurement of cutting resistance and cutting temperature
Difference of machinability are often examined by analyzing cutting force or cutting temperature.
However since in 3-dimensional cutting, cutting force analysis is very complicated and cutting temperature
measurement is extremely difficult for end-mill cutting in machining center, cutting resistance and cutting
temperature were measured by two dimensional turning. As shown in Figure 1, by grooving turning test
pieces with 65 mm diameter, disk parts with 2 mm in thickness were produced. Table 3 shows the conditionof two dimensional turning test. Piezoelectric cutting dynamometer was used to measure cutting resistance.
Principal partial force and feed partial force are measurement items and machinability indicators. Cutting
mechanism were considered by analyzing cutting forces. Cutting temperature were measured by tool-work
thermocouple method as shown in Fig. 1. Thermoelectromotive force generated during turning wasmeasured and temperature was calculated from calibration curve produced by data previously measured
using same material combination. Cutting temperature measured by this method are said to be mean
temperature between chip-tool contact area10).

Fig. 1. Measurement method of cutting force and cutting temperature.

© 2020 Daido Steel Co., Ltd.

No. SD2005 (2020.08)

4/17
Table 3. Cutting conditions for cutting force and cutting temperature measurement in turning
Tool grade

M20(Carbide), Non coating

Tool geometry

TNGA160404

Cutting speed

50 m/minFeed

0.15 mm/rev.

Width of cut

2 mmCutting fluidDry(4) Microstructure observation of chips
It is known that cutting shape was reflected by cutting state difference11). Therefore observation of
longitudinal section of chips generated by turning were taken place. After polishing observation face,etching was done by Nital before optical microscopic observation. SEM back scattered image observationof as-polished samples were also taken place and carbide precipitation status was investigated. Moreover inorder to confirm oxidation status of chip surface, Transmission Electron Microscope (TEM) observation
for sections of chip surface was taken place. Film specimen were taken from chips generated by turning andwere prepared using Focused Ion Beam Spattering (FIB) as shown in Figure 2.

Fig. 2. FIB method to prepare cross-sectional specimens for TEM.
(5) High temperature high speed tensile test
It is known that cutting chips are formed by extremely large plastic deformation in primary shearing stressarea and secondary shearing stress area and their strain speed is likely to be in the order of 102s – 105/s.
Therefore Gleeble tensile equipment which enables high speed tensile testing was utilized to perform hightemperature tensile test in order to investigate cutting mechanism from high temperature deformation
morphology. Geometry of tensile test pieces are 4.5mm in diameter and 25 mm long of parallel part. They
were taken from billets so that their axis direction is as same as forging direction. Tensile testing was
performed with the speed of 800 mm/s after the temperature was raised to 873 – 1273 K in 100 s and was
kept at the testing temperature for 60 s, and deformation load and reduction of area at fracture were
measured. The strain speed in this condition was calculated as about 50/s. This value is larger than normal
static tensile test, but is smaller than actual cutting deformation.

© 2020 Daido Steel Co., Ltd.

No. SD2005 (2020.08)

5/17

3. Results and discussion
3.1 Microstructure observation
Optical microscopic observation results of hardened and tempered structure of each sample are in Figure
3. They were all martensitic structures and showed no large differences by Si content.

Fig. 3. Microstructures of test specimens.
Figure 4 shows SEM observation results of as-hardened and hardened and tempered steel A, C and E. In
SEM observations, back scattered images were used to clarify the differences in carbide compositions.
White colored areas and black colored areas were analyzed by Energy Dispersive X-ray Spectroscopy (EDX)
and concentration of Mo in white areas and that of V in black area were confirmed, which lead to M6C
carbides and MC carbides respectively. Sizes of M6C and MC carbides are about 0.5 μm and the amountof these carbides decreases drastically with decrease of Si content. It is noted that since no large difference
was confirmed in the amount of carbides with or without tempering, very fine carbides which precipitate
during tempering were not captured by SEM observations. Therefore observed carbides in this study are
considered to be residual carbides which had existed during forging and unsolved during hardening.

© 2020 Daido Steel Co., Ltd.

No. SD2005 (2020.08)

6/17

Fig. 4. Back scattering electron images of carbides in hardened specimens and tempered ones.
Now stability of carbides were calculated using thermodynamic equilibrium software, Thermo-Calc.
Figure 5 shows equilibrium phase diagram of target composition system with change of Si content. This
phase diagram shows decrease of solution temperature of M6C and MC carbides with decrease of Si content,
which explains decrease of residual carbides in low Si steel qualitatively. For example, M6C carbides do notexist in steel with less than 1% Si content while they do exist in steel with more than 1% Si. Actually not
many M6C carbides were confirmed in low Si steel, thus observation results and calculated phase diagramby Thermo-Calc were matching well. However since production process is not carried out underequilibrium condition, some M6C carbides remain insignificantly.
From above results, Si has the effect of raising solution temperature of M6C and MC carbides and withdecrease of Si content the solution temperature of M6C and MC carbides decrease resulting in decrementof carbide amount which remain at hardening treatment.

Fig. 5. Vertical section phase diagram of theFig. 6. Effect of Si content on tool life in

Fe-0.35%C-Si-0.6%Mn-5.5%Cr-3%Mo-

end-milling.

0.85%V system calculated by thermo-calc.

© 2020 Daido Steel Co., Ltd.

No. SD2005 (2020.08)

7/17

3.2 Machinability of end-milling
Figure 6 shows the relationship between Si content and tool life in end-milling. Tool life declines withdecrease of Si content and the amount of tool life decrease tends to larger in the region where Si content isbelow 0.3%. It is clarified that, as it has been reported for conventional relatively low hardness material4), 5),
machinability declines with the decrease amount of Si for hardened and tempered hot working die steelswith hardness of 49 HRC which is the maximum limit of usage. Moreover, comparing with past reports on
the view point of large drop of tool life, in which tool lives decline drastically around less than 0.3 to 0.5%
Si4), 5), this study’s result is almost the same.
Figure 7 shows SEM observation results of damaging process of flank face of carbide tools. The lower row
of each photo corresponds to characteristic X-ray image of Fe, which explains adhesion of test material.
Steel A having low Si content, shows adhesion of material at the cutting length of 200 mm and the adhesionarea becomes larger and detachment can be observed at the cutting length of 400 mm. On the other handsteel D having high Si content does not show any adhesion of material at the cutting length of 400 mm andat the cutting length of 800 mm adhesion and detachment can be observed. Therefore the cause of decline
in tool lives for low Si material can be considered that the adhesion of material occurs more easily with low
Si steel than high Si steel, adhesive wear progresses more rapidly. Besides, although for steel B as well as
steel A adhesion of material was observed at the cutting length of 200 mm, for steel C adhesion was notobserved at 200 mm length but was observed at 400 mm length. The reason why tool lives drop drasticallybelow 0.3% Si steel is considered to acceleration of adhesion starting timing occurs below around 0.3% Si.

© 2020 Daido Steel Co., Ltd.

No. SD2005 (2020.08)

8/17

Fig. 7. SEM and characteristic X-ray images of clearance face of carbide tool in end-milling.
3. 3 Results of cutting resistance and cutting temperature measurement
Figure 8 shows the relationship between Si content and cutting resistance in case of turning. All principal
partial force, Fp, feeding partial force, Ff and cutting resultant force, R tend to increase with decreasing Si
content, so machinability in turning process is supposed to decrease with Si content decrease. The analysis
using these results will be discussed in the last section.
Figure 9 shows relationship between Si content and cutting temperature. Cutting temperature reached
915 to 1165 K and increases with decrease in Si content. Temperature difference between steel A and steel
E is 250 K, therefore influence of Si on cutting temperature is considered to be significant. Moreover, thecolor of chips turned to light blue from purple in accordance with decrease of Si content, which proves thatthe cutting temperature for low Si steel is high12). From this result, it is assumed that the cutting temperaturetends to rise easily for low Si steel under same cutting conditions. In case of end milling, although cutting
mechanism is different, it is also assumed that the cutting temperature is high for low Si steel and adhesive
ware tends to occur more easily. AC1 transformation temperature of each test material is about 1100 K, and© 2020 Daido Steel Co., Ltd.

No. SD2005 (2020.08)

9/17
the cutting temperature of steels with Si content lower than 0.3% seems to be higher than AC1 temperature
as shown in the phase diagram (Fig. 5). Even though it was not verified yet, adhesion of materials is related
to AC1 transformation temperature. It is possible to assume that tool life drops drastically below certain Si
content because the cutting temperature overtakes AC1 transformation temperature.

Fig. 8. Effect of Si content on cutting forces inFig. 9. Effect of Si content on cutting

turning.

temperature in turning.

3. 4 Microstructure observation of chips
Figure 10 shows optical microscopic observation results for chips generated by turning, and upper side of
pictures correspond with back sides of chips (contacting sides to tool’s rake face). Chip shape is all flowtype11), and the plastic flow by primary shear deformation can be recognized. Moreover, white layers of
about 10 μm in thickness on surfaces of back sides of chips are formed, which might be caused by thermal
effect. Chip shapes and microstructures did not show large differences depending on Si content, therefore
cutting states did not seem to be very different.
SEM observation results of chip surfaces are shown in the right column of Fig. 10. They were backscattered images of back sides of chips and the solution of M6C carbides can be observed in the area below
about 10 μm from the surface. This means that these areas were heated above 1303 K which corresponds
to their hardening temperature, and surfaces of back sides of chips can be considered to be heated above
1303 K by friction heat. It is also assumed that, even if it is momentary phenomenon, outermost layer area
must be heated to several hundred degree more than average chip-tool contact face temperature. Carbide
solution state depends on Si content, and M6C carbides were hardly observed for steel A – C. As describedearlier, this phenomenon can be explained by the fact that solution temperature of carbides are low and thecutting temperature is high in case of low Si steels. White layers which were observed by optical microscopy
can be assumed to be transformed structure which was quenched as the result of heated to high temperatureand cooled by heat diffusion.

© 2020 Daido Steel Co., Ltd.

No. SD2005 (2020.08)

10/17

Fig. 10. Microstructure and back scattering electron images of chips in turning.
Okada et al. investigated the relationship between machinability and Si content, by focusing on
compositions of oxides which were formed on surfaces of chips. In their report that machinability of 0.3C4Cr-1.5Mo-0.6V-Ni steel with 43 HRC hardness decrease with decrease of Si content, the difference inmachinability is caused by difference in chip deposition property by oxides composition change after
performing secondary ion mass spectroscope analysis, finding out condensation of Cr in oxides of low Si(0.1% Si) steel and their oxides consist of mainly Fe and Cr oxides, while condensation of Cr were notobserved in oxides of high Si (1.6% Si) steel and their oxides consist of mainly Fe and Si oxides5). Moreover,
it is reported that the latter oxides have low melting point and lubrication between chip and tool wereenhanced13). Therefore chips of steel A and D generated in turning were analyzed after making film sample
by FIB method and analyzing oxides compositions by TEM-EDX. Figure 11 shows the result of TEM image
and EDX analysis. Both photos show oxide layers of about 0.05μm on their surfaces, but no condensationof Cr or Si were observed. Quantitative analysis was performed and the ratio of Fe, Cr and Si in the oxide
was nearly the same as processed material. Therefore although oxides formed on low Si steels are mainly
consist of Fe and Cr, those formed on high Si steels are not always consist of mainly Fe and Si, andcontaining the same level of Cr as low Si steels. The reason why our result is different from theirs is notclear, but cutting method difference can be considered. The difference in lubrication effect will be discussed
in the last section.

© 2020 Daido Steel Co., Ltd.

No. SD2005 (2020.08)

11/17

Fig. 11. TEM images of back face of chips and depth profiles of Fe, Si, Cr and O by EDX analysis.
3. 5 High temperature tensile test characteristics
Figure 12 shows the relationship between Si content and maximum deformation stress and Figure 13
shows the relationship between Si content and reduction of areas. Changes in maximum deformation stressand reduction of area are small in 1073 – 1273 K region. This is considered that the temperature is abovetransformation and composition dependence is small. On the other hand, both maximum deformationstress and reduction of area increase with decrease of Si content at 873 and 973 K. As shown in Fig. 4,amount of residual carbides is smaller in low Si steel, thus the mechanism of change in high temperaturetensile test characteristics depending on Si content can be considered as follows. Increase of maximum
deformation stress with decrease of Si content is caused by amount increase of fine carbides which
precipitate during tempering or by amount increase of alloying elements which dissolve in matrix phase
after tempering. However it is not clear which of these elements, precipitation hardening or solution
hardening is dominant to high temperature strength. This will be a future assignment. In ductile fracture,
voids nucleated by secondary particles in materials grow and combine until completion of necking process14).
Therefore increase of high temperature ductility was due to decrease of retained carbides which can be
nuclei of voids.
Moreover, it is known that temperature in primary shear stress area where chips are generated is lower
than cutting temperature15). Temperatures in shear area of turning process in this study are supposed to belower than 973 K according to Fig. 9. This temperature range is where tensile characteristics changes
depending on Si content. These tensile characteristics differences seems to be affect machinability. This
will be discussed in the next section.

© 2020 Daido Steel Co., Ltd.

No. SD2005 (2020.08)

12/17

Fig. 12. Effect of Si content on maximum stressFig. 13. Effect of Si content on reduction of

in high speed tensile test at elevatedarea in high speed tensile test at elevatedtemperatures.

temperatures.

3. 6 Mechanism of machinability transition
As a result of these investigation, these facts that machinability of hot working die steels in high hardness
state will deteriorate with decrease of Si content and high temperature strength and hot ductility are betterfor low Si steels were found. Hereinafter mechanism of machinability change with Si content was examinedusing results of cutting resistance measurement of turning, calculating each cutting partial force by singleshear plane model16).

Fig. 14. Two-dimensional cutting model.
Friction force, F and vertical force, N on the rake face of tools occur as shown in Figure 14, and tool rake
angle is described as α then following equations can be derived.
F = Fp sin 𝛼 + Ff cos 𝛼 ⋯ ⋯ (1)N = Fp cos 𝛼 − Ff sin 𝛼 ⋯ ⋯ (2)On the other hand, apparent friction coefficient, μ and cutting shear angle, Φ can be described as© 2020 Daido Steel Co., Ltd.

No. SD2005 (2020.08)

13/17
following equations.
μ = tan 𝛽 = F/N ⋯ ⋯ (3)tan 𝜙 = rC cos 𝛼 /(1 − rC sin 𝛼) ⋯ ⋯ (4)Here β shows friction angle, rC represents cutting ratio and described as following definition.
rC = t/t C ⋯ ⋯(5)
Now focusing on cutting shear faces, cutting resultant force, R can be resolved into shear direction partial
force, FS and vertical force, NS, while both factors can be described as follows.
FS = 𝐹𝑃 cos 𝜙 − 𝐹𝑓 sin 𝜙 ⋯ ⋯ (6)NS = 𝐹𝑃 sin 𝜙 + 𝐹𝑓 cos 𝜙 ⋯ ⋯ (7)Shear stress, τS can be described using cutting width, b.
τS = FS sin 𝜙 /(t ∙ b) ⋯ ⋯ (8)Moreover, from above equations friction force on rake faces, N and vertical force, N can be derived asfollows.
F = 𝜏𝑆 ∙ 𝑡 ∙ 𝑏 sin 𝛽 / sin 𝜙 cos(𝜙 + 𝛽 − 𝛼) ⋯ ⋯ (9)N = 𝜏𝑆 ∙ 𝑡 ∙ 𝑏 cos 𝛽 / sin 𝜙 cos(𝜙 + 𝛽 − 𝛼) ⋯ ⋯ (10)In cutting equations that describe the relation between shear face shear angle,Φ and rake faces friction
angle,β, if substitution is made with Lee-shaffer’s equation (11) that showed good agreements with thisstudy’s results,
𝜙 = π/4 − β + α ⋯ ⋯ (11)following equations can be acquired.
N = √2 ∙ 𝜏𝑆 ∙ t ∙ b cos 𝛽 / sin(𝜋/4 − 𝛽 + 𝛼) ⋯ ⋯ (12)F = √2 ∙ 𝜏𝑆 ∙ t ∙ b sin 𝛽 / sin(𝜋/4 − 𝛽 + 𝛼) ⋯ ⋯ (13)Equations (12) and (13) indicate that if t, b and α are known values, by determining either τS or β,
cutting forces, F or N can be estimated as functions of τS or β.
Based on above discussions, Figure 15 shows friction forces, F and vertical forces, N on rake faces of eachtest material calculated from eq. (1) and (2), as well as change of cutting force when only τS decreases
keeping β constant as a solid line arrow, and when only friction angle β decreases keeping τS constantas broken line arrow. The change of cutting force when Si content is increased compared to steel A isassumed to be close to the solid line arrow. This suggests the change of cutting forces is not brought about
by decrease of friction angle or friction coefficient, but by decrease of shear stress.

© 2020 Daido Steel Co., Ltd.

No. SD2005 (2020.08)

14/17

Fig. 15. Influence of change in shear stressFig. 16. Effect of Si content on shear stressand friction angle on normal force andacting on shear plane in turning.

friction force in turning.
Therefore, the relationship between Si content and shear stress were calculated using eq. (8). Figure 16
shows those results. Increase of shear stress can be confirmed with decrease of Si content, drastic increase
can be seen especially in the region where Si content is less than 0.3%. From above discussion the cutting
resistance increase with decrease of Si content con be considered as increase of shear stress. Increase ofshear stress originates from high temperature strength judging from the high temperature tensile test result
even though strain speed is different. There is another theory that decrease of shear stress comes fromhigher cutting temperatures for low Si steels. However since temperatures in shearing area are assumed to
be lower than 973 K as described previously, shear stress change should be small even if temperatures inshearing area change in case of cutting deformation where strain speed is very high17). On the other note,the most effective factor on the cutting temperature is known to be shear stresses as M. C. Shaw’s analysis
on the cutting temperature18). Raise in the cutting temperature or acceleration of adhesion wear along withdecrease of Si content are also supposed to be the one of causes of increase of high temperature strength.
Moreover, changes of friction angles were investigated. Nakamura et al. have investigated the relationshipbetween cutting resistances and friction angles using JIS SCM420 grade steels with additions ofmachinability enhancing elements19). They have reported in the paper that the decrement of ductility inhigh temperature high speed tensile testing had decreased friction angles and cutting frictions. Therefore,
fracture strain,εf was derived from reduction of area in high temperature high speed tensile testing, ψ,
while friction angle β was computed from eq. (3) in order to examine the relationship between them using
eq. (14).
εf = ln(1/(1 − 𝜓)) ⋯ ⋯ (14)© 2020 Daido Steel Co., Ltd.

No. SD2005 (2020.08)

15/17

Fig. 17. Relationship between fracture strain in tensile test and friction angle in turning.
Figure 17 shows the relationship between fracture strains and friction angles. Data by Nakamura et al. wereadded to this figure. Overall tendency, where friction angles decrease with decrement of fracture strains
and decrease of high temperature ductility reduces friction angle, shows similarity. However compared withdata of Nakamura et al. where they studied effect of machinability enhancing element, the amount ofdecrease in friction angles depending on the decrease of fracture strains is smaller in this study. This may
be the result of small lubricant effect between chip-tool by inclusions. Therefore in materials dealt in thisstudy changes of friction angles are relatively small even though Si content was reduced and the influenceof friction angle in other words friction coefficient to the cutting force is considered to be small. This result
supports investigation results evaluated in Fig. 15. Even though the lubrication effect increases depending
on compositions of oxides, that effect is thought to be small. Lastly, summary of this examination is
illustrated by changes of factors related to cutting processes and their relationships in Figure 18. Under theassumption where the result of turning processes can be applied to end-milling processes, the reason why
the tool life decreases with decrease of Si content can be considered as follows.
Si content decrease↓
Increase of high temperature strength↓
Increase of shear stress↓
Increase of cutting resistance, increase of cutting temperature↓
Acceleration of adhesion wear
However even though influences are thought to be rather small, changes in high temperature ductility or
friction angles were confirmed. It will be necessary to investigate effects of high temperature ductility or
friction angles by examining under different hardness levels or different machining conditions.

© 2020 Daido Steel Co., Ltd.

No. SD2005 (2020.08)

16/17

Fig. 18. Changes of factors on machining with decreasing Si content.

4. Conclusion
This study has been carried out to clarify the effect of Si content on machinability of 5%Cr-3%Mo type
hot working die steel hardened and tempered to 49 HRC. The results are as follows.
(1) Carbides retained in hardened and tempered specimens are M6C and MC and the amount of thesecarbides decrease with the decrease in Si content.
(2) In end-milling, carbide tools exhibit shorter life with decreasing Si content, especially at the range ofSi content less than 0.3%. Adhesive wear increases with the decrease in Si content.
(3) In turning by carbide tools, the cutting forces and temperatures rise with decreasing Si content.
(4) High speed tensile test at elevated temperatures reveals that the steels with lower Si content show
higher maximum flow stress and larger reduction of area.
(5) By the calculation of turning-cutting forces based on the two-dimensional cutting model, it is foundthat the decrease in Si content raises the shear stress resulting in the increase of cutting forces. The highshear stress of lower Si steels is considered due to their high strength at elevated temperature.

5. Acknowledgements
Great gratitude has to be expressed to Mr. Masahiro Kawasaki and Mr. Toshiaki Suzuki from Nippon
Denshi, who cooperated with us by making film specimens using FIB method and TEM observations.
(Reference)
1) K. Sudou: Denki-Seiko (Electric Furnace Steel), 60 (1989), p. 367
2) T. Okuno: Tokushukou (The Special Steel), 46 (1997), No. 7, p. 6

© 2020 Daido Steel Co., Ltd.

No. SD2005 (2020.08)

17/173) T. Ogawa: Tokushukou (The Special Steel), 46 (1997), No.7, p. 23
4) K. Iwamoto: Denki-Seiko (Electric Furnace Steel), 64 (1993), p. 191
5) T. Sera et al.: CAMP-ISIJ, 9 (1996), p. 1404
6) M. Fruya et al.: Kikai Gijutu, 44 (1996), No. 8, p. 39
7) E. Takeoka: Kikai Gijutu, 45 (1997), No. 10, p. 228) T. Nishimura: Netushori (Journal of the Japan Society for Heat Treatment), 20(1980), p. 342
9) T. Sugita et al.: Kiso-Sessaku-Kakougaku, (1984), p. 87 [Kyoritu Shuppan]
10) Y. Fujimura: Jitsuyou-Sessaku-Kakouhou, (1989), p. 50 [Kyoritu Shuppan]
11) E. Shirai: Sessaku-Kensaku-Kakouhou, Part I, (1971), p. 8 [Kyoritu Shuppan]
12) Y. Fujimura: Jitsuyou-Sessaku-Kakouhou, (1989), p. 170 [Kyoritu Shuppan]
13) Y. Okada: Netushori (Journal of the Japan Society for Heat Treatment), 39 (1999), p. 225
14) Zairyou-Kyoudogaku, (1986), p. 19 [Nihon Zairyou Gakkai]15) Y. Lee et al.: Seimitukikai, 44 (1978), p. 1424
16) E. Shirai: Sessaku-Kensaku-Kakouhou, Part I, (1971), p. 16 [Kyoritu Shuppan]17) K. Nakayama: Sessakune-Kakouron, (1978), p. 109 [Korona Sha]18) E. G. Loewen, M. C. Shaw: Trans. AMSE, 76 (1954), p. 217
19) S. Abeyama et al.: Denki-Seiko (Electric Furnace Steel), 55 (1984), p. 177© 2020 Daido Steel Co., Ltd.

No. SD2005 (2020.08)