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Tribological behaviour and statistical experimental design of sintered iron–copper based composites
Applied Surface Science 285P (2013) 72–85
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Tribological behaviour and statistical experimental design of sintered iron–copper based composites Ileana Nicoleta Popescu a , Constantin Ghit¸a˘ b , Vasile Bratu a,∗ , Guillermo Palacios Navarro c a b c
Valahia University of Targoviste, Faculty of Materials and Mechanics Engineering, 18-24 Unirii Boulevard, 130082, Targoviste, Romania Valahia University of Targoviste, Faculty of Science and Arts, 18-24 Unirii Boulevard, 130082, Targoviste, Romania University of Zaragoza, Department of Electronic Engineering and Communications, Zaragoza, Spain
a r t i c l e
i n f o
Article history: Received 12 April 2013 Received in revised form 30 July 2013 Accepted 1 August 2013 Available online 11 August 2013 Keywords: Tribological characteristics ESEM EDS X-ray maps and compo image of surface friction materials Statistical experimental design
a b s t r a c t The sintered iron–copper based composites for automotive brake pads have a complex composite composition and should have good physical, mechanical and tribological characteristics. In this paper, we obtained frictional composites by Powder Metallurgy (P/M) technique and we have characterized them by microstructural and tribological point of view. The morphology of raw powders was determined by SEM and the surfaces of obtained sintered friction materials were analyzed by ESEM, EDS elemental and compo-images analyses. One lot of samples were tested on a “pin-on-disc” type wear machine under dry sliding conditions, at applied load between 3.5 and 11.5 × 10−1 MPa and 12.5 and 16.9 m/s relative speed in braking point at constant temperature. The other lot of samples were tested on an inertial test stand according to a methodology simulating the real conditions of dry friction, at a contact pressure of 2.5–3 MPa, at 300–1200 rpm. The most important characteristics required for sintered friction materials are high and stable friction coefficient during breaking and also, for high durability in service, must have: low wear, high corrosion resistance, high thermal conductivity, mechanical resistance and thermal stability at elevated temperature. Because of the tribological characteristics importance (wear rate and friction coefficient) of sintered iron–copper based composites, we predicted the tribological behaviour through statistical analysis. For the first lot of samples, the response variables Yi (represented by the wear rate and friction coefficient) have been correlated with x1 and x2 (the code value of applied load and relative speed in braking points, respectively) using a linear factorial design approach. We obtained brake friction materials with improved wear resistance characteristics and high and stable friction coefficients. It has been shown, through experimental data and obtained linear regression equations, that the sintered composites wear rate increases with increasing applied load and relative speed, but in the same conditions, the frictional coefficients slowly decrease. © 2013 Elsevier B.V. All rights reserved.
1. Introduction The increasing consumer demand for low-cost performing and non-polluting materials, with a favourable price/quality ratio, has determined, in the past decades, development of extensive researches in automotive brake friction industry. The efforts include the development of ferrous (composites with iron based matrix) and non-ferrous based materials such as copper, aluminium or carbon composites, as new candidates of brake friction materials [1–6]. Friction materials are used for converting the kinetic energy of the automotive to thermal energy (heat), resulting from two surfaces in contact of the brake pads and brake disc, which is absorbed and dissipated by the materials [7].
∗ Corresponding author. Tel.: +40 245206106. E-mail address: vasilebratu.uvt [email protected] (V. Bratu). 0169-4332/$ – see front matter © 2013 Elsevier B.V. All rights reserved. http://dx.doi.org/10.1016/j.apsusc.2013.08.007
The brake friction materials have a complex composition that includes both metallic and non-metallic components. Also, brake pads typically comprise the binders, the lubricating components and frictional additives [7]. The purpose of a binder is to maintain the brake pads’ structural integrity (to hold the components of a brake pad together) under mechanical and thermal stresses. The binders could be synthetic resins, metallic ingredients or mixture of metallic and organic ingredients. The synthetic resin based friction materials is cheap to produce, but it has as a disadvantage that, in high-energy braking applications (operating up to 370–450 ◦ C) when the induced temperature can be high enough, the material decomposes or breaks down. As a result, the friction coefficient with the brake disc is compromised leading to decreasing road safety [7,8]. In case of using metallic binders as iron, cooper, aluminium and their alloys, we can improve the wear resistance, thermal diffusivity and strength of the brake friction material. The lubricating components such as graphite, molybdenum disulphide, etc. have the objective to: (i) stabilize the developed friction coefficient
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during braking, particularly at high temperatures, (ii) decrease wear of counterparts and also (iii) increase grippe resistance [1–14]. The frictional additives, such as oxides (Al2 O3 , SiO2 , ZrO2 , Mullite, and Spinel), nitrides (TiN and Si3 N4 ), carbides, (SiC, TiC, B4 C, and VC), determine the frictional properties of the brake pads. Also the frictional components increase and stabilize the friction coefficient values, respectively increase the wear resistance [1–12,15]. Composites with iron or steel matrix and ceramic reinforcements used like friction materials bring new possibilities in the production of wear resistant materials because they provide good wear resistance and maintain friction effectiveness at elevated temperature [7,11,15]. Iron could be alloyed with Cu, Cu–Sn–(Zn), Ni, Mo, Al, Co, Mn, etc. for improving resistance of the metallic matrix [7–11,15]. Cooper and Cu alloys are chosen mainly to improve the thermal conductibility at the friction interface and for sustaining the level of the coefficient of friction at elevated temperatures by producing copper oxides at the friction interface [1,8]. Aluminium or aluminium alloys are used in frictional materials because of lightweight and good thermal diffusivity [12,16–18]. The proportion and granulation of these components, was and is still the subject of numerous investigations in various countries in order to achieve the optimal characteristics of brake friction materials. The processing of brake friction materials is realized by a diversity of methods [6–8,13–23] including powder metallurgy (P/M) techniques. The P/M techniques give (i) a good uniformity distribution of lubricated and frictional components in metallic matrix, (ii) near net-shape parts fabrication, (iii) high productivity (iv) dimensional accuracy and (v) allows a good control of brake friction require characteristics. Among the required market performance criteria for brake pads materials, we mention [8]: (a) maintaining a high and stable friction coefficient with the brake disc, in various conditions, including at high temperatures; (b) very good wear resistant material and implicit very good durability in service; (c) good strength at elevated temperatures; (d) high thermal conductivity; (e) high corrosion resistance; (f) smooth braking assurance. In this paper we obtained frictional composites by conventional P/M technique and we have characterized them by microstructural and tribological point of view.
2. Materials and compositional selection The raw powders chosen for frictional materials were Fe (reduced), Cu (electrolytic) as metallic ingredients, graphite and MoS2 as solid lubricant, respectively, mullite and spinel as frictional components, with a particles size situated up to 100 m. The raw powders selected must be non-polluting material from compositional point of view and have required characteristics. Using iron/steel based material in comparison with Cu-based material improved strength, hardness, ductility and heat resistance properties. Iron as a matrix of the automotive brake material has been chosen due of its stability under higher temperatures (1100 ◦ C), which imply very good tribological behaviour under heavy duty dry operating conditions. The selection of copper as alloying element gives additional strength of sintered steels (the elemental copper at 1083 ◦ C promotes sintering and enhances the strength of the steel) [10]. In order to improve cold and hot strength of sintered friction materials a tight ratio between cooper and iron was chosen. Generally for the friction sintered materials for automotive brakes, the cooper ratio in the material is majority [3]. Graphite + MoS2 have the role to stabilize the developed friction coefficient during braking, particularly at elevated temperatures. Additions of graphite to Fe–Cu sintered steels are desirable because the carbon promotes the formation of a pearlitic microstructure, resulting in additional strength and hardness of steel [9].
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The high proportion in lubricating components was chosen due to (i) easy pressing of the mixture, at medium pressure of 500–600 MPa; (ii) improvement of sliding properties, by forming a lubricant layer on the opposing counter friction material at speed of 20–30 m/s [3,7,14]. The new chemical compositions (new proportion of components) established after a number of preliminary tests [15] of iron–copper based friction material (SFM1 and SFM2) samples preparation and the chemical composition of a commercial brake pad, used as reference material (RM) are presented in Table 1. 3. Materials and experimental procedure 3.1. Powder analysis The following characteristics of powders were analyzed: (i) the apparent density according to SR EN 23923-1: 1998 standards, (ii) flow rate, in concordance with the standard method SR EN ISO 3953: 1998; (iii) particle size distribution, according with SR 1320394 and ISO 4497:1983; (iv) the morphology of powder particles determined by scanning electronic microscopy (SEM). 3.2. Composite preparation We used P/M as a technique for preparation of composites for automotive brake pads because this technology allows us to obtain sintered friction materials with a large variety of chemistry in concordance with different conditions of solicitation. Before the manufacturing of the composite, the oxides of iron copper powders were reduced in “Siemens–Plania” type furnace in presence of H2 atmosphere at temperature near 980 ◦ C for iron and 280 ◦ C for copper, during 60 min, in order to obtain reduced metallic powders. The ceramic powders were heated for 300–400 ◦ C, holding 120 min, for elimination of adsorbed gases, moisture and organic contaminants. The dosage of mixtures was made gravimetric and the elemental powder of the mixtures was dry blended using the Double Cone Blender (10 kg capacity) at a rotation speed of 20 rpm, during 6–8 h. The mixed homogenous powders were compacted at room temperature in a double action hardened steel die with a automate hydraulic press of 30 tone force, Meyer Type, at pressure of 500–600 MPa. According to literature [23–25] there are a diversity of process parameters (sintering temperature, sintering time, and protective environment). For our experiments, the compacts were sintered in a sintering furnace, Balzers type, at 1050 ◦ C sintering temperature, 90 min holding time, in vacuum atmosphere under a pressure of 0.1–0.3 mmHg, in limits mentioned of scientific literature. The composite processing flow-sheet is presented schematically in Fig. 1. 3.3. Sintered friction material characterization The obtained sintered SFM1, SFM2, composites and reference material RM were analyzed from physical, mechanical, microstructural (SEM) and micro-compositional (SEM, EDS and X-ray maps) point of view, according with specific techniques presented in [26]. The quantitative X-ray microanalysis of obtained sintered iron–copper based composite as frictional material was carried out using the energy dispersive X-ray spectroscopy (EDS) of the environmental scanning electron microscopy (ESEM), FEI XL-30 and FEG SEM Inspect F, FEI. The density of the sintered composites was measured by physical measurements and also using the Archimedes method. The Brinell Hardness test was done according ISO 4498/1 standardization.
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Table 1 Raw powders chemical composition (wt.%) of brake friction materials. Friction material type
SFM1 SFM2 RM
Metallic binder
Lubricant component
Frictional components
Fe
Cu
Sn
Graphite
MoS2
Mullite (3Al2 O3 ·2SiO2 )
24–30 30–35 30–32.5
21–25 20–30 30–32.5
– – 2–3
18–22 13–17 9–11
4–8 5–7 4–6
The tribological (friction and wear) tests of sintered iron–copper based composite for automotive brake pads were performed using a “pin-on-disc” tribometers. One lot was tested with a tribometer (with a grey cast iron of 300 HB as counterparts), on 10 mm × 10 mm × 7 mm samples, in dry friction conditions, applied normal load between 0.35 and 1.15 MPa, during 15–20 min per sample, at 12.5–16.9 m/s relative speed in braking points (sliding distance between 8125 and 14733 m) and at constant temperature. The test machine consists of a balanced pendulum in which is mounted the sample into a box. The sample rubs on a cast iron disc and the disc is acted through the main body of a transmission belt, by an electric motor of 1.4 kW power. The peripheral speed of the disc can be varied by positioning the sample box (sample position) at different distances from the axis of the machine through a variable ratio transmission. The frictional force was determined by measuring the angular movement (˛) of pendulum from the vertical position (Fig. 2). The correlation between applied pendulum weight (G) in Kg, total length (L) in mm and the distance from the spindle axis to pendulum test box (e) in mm is given by the relation. F = L · G · sin ˛/e
(1)
Spinel (MgAl2 O4 )
Balanced Balanced Balanced
−
From Eqs. (1) and (2) result that the coefficient of friction is determined by the relationship: = L · G · sin ˛/Ne
(3)
The wear rate was measured, respectively, by height loss (mm/h) and weight loss (g/cm2 h). The wear of the friction materials was expressed by weight/height loss percent as: W % = (W0 − W1 )/W0 · 100%
(4)
where W0 and W1 are the weight/height of the specimens before and after testing, respectively. The other lot of samples were tested on an inertial test stand according with Fermit S.A. methodology [15]. In this type of testing the real conditions of dry friction of automotive brake pads were simulated (cold and hot testing, grinding and endurance conditions) on dry sliding wear conditions, at a contact pressure of 2.5–3 MPa, depending on braking parameters as: (i) number of braking stops (25–500) and (ii) speed 300–1200 rpm. The temperature during braking ranged between 75 and 350 ◦ C. The dimensions of sintered friction samples of the second lot, for test samples on inertial stand, fixed on a metallic pad, were 2500 mm2 × 12 mm (Fig. 3).
It is known that: = F/N
(2)
where F is the frictional forces; is the coefficient of friction and N is the normal forces.
Fig. 1. The P/M technology route of SFM1 and SFM2 materials for braking pads.
3.4. Factorial design Because of the tribological characteristics importance (wear rate and friction coefficient) of sintered iron–copper based composites, we predicted the tribological behaviour through statistical analysis [27–30]. For that purpose, a factorial design of experiment of the type Pn [30] was used in the present paper, for the testing sintered frictional material from the first lot, in dry friction conditions. The “n” corresponds to the number of factors and “P” stands for the number of levels.
Fig. 2. Schematic diagram of the forces which activate during the wearing of samples (pin-on-disc).
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Table 3 The average values concerning physical and mechanical characteristics of iron–copper sintered composites for automotive brakes. Sintered friction material
Density, g/cm3
Hardness, HB
SFM1 SFM2 RM
3.8 3.9 3.6
16.5 19.2 11.8
4. Results and discussion
graphite and MoS2 have the particles shape as flaws, especially the molybdenum disulphide particles have the tendency of coalescence. The average values of apparent densities of the mixtures for frictional materials were 1.39 g/cm3 for SFM1 powder mixture and 1.46 g/cm3 for SFM2 mixture. The physical and mechanical characteristics (Brinell hardness and density) of the obtained composite friction material, in comparison with reference material, are presented in Table 3. The structure of iron–cooper composites, the ESEM, EDS analysis and X-ray mapping of the elements from SFM1 samples are presented in Figs. 5–8. The analyses were made on the surface of the frictional composite material, from different white and grey compact areas. In Figs. 5 and 6 we can observe the components and elements from the surface of sintered iron–copper based composite. The microstructure and elemental analysis shows a basic mass formed by iron and copper (white compact areas) as we observed in Fig. 5(a) and (c), respectively and in Fig. 7(a) and (b). The mechanical inclusions, especially mullite, spinel, are identified from X-ray mapping (Al, O, Si, and Mg elements) and EDS elemental analysis (the proportion of them), in Fig. 6, and we observed them from microstructural point of view as light grey compact particles, Fig. 7(a) and (d), respectively and Fig. 8. We can notice the graphite (C) as dark grey compact particles in all ESEM microstructure and confirmed by X-ray mapping (Ck) and EDS elemental analysis. Due to the combined influences of powder mixtures components, it is not possible to determine precisely the influence of each component on the basic structure. The most pregnant structural modifications are determined especially by the copper additives. The equilibrium diagram of Fe–Cu shows that copper is liquid at 1083 ◦ C and can easily infiltrate between the Fe particles. There is a partial diffusion in iron and, thus, the net is reinforced (liquid copper ensures a better connection between alloy particles and resulted that by action of capillary force). The above described phenomena can take place at lower temperatures than the melting points of copper, due to the iron-solubility of 8.5% at 1094 ◦ C and of 4.5% at 833 ◦ C.
4.1. Powders and frictional material characteristics
4.2. Influence of braking parameters on friction coefficients
In Table 2 the physical characteristics of raw powders are presented and in Fig. 4 the morphology of elemental powders is presented. We observed in Fig. 4 that iron powders have spherical and egg-shape morphology, the dendrites particle shape is specific for electrolytic copper powders, the mullite has irregular form and the
The braking parameters used were number of braking stops (25–500) and speed of 300–1200 rpm. The friction coefficient variation depending on temperature, the speed and number of brakes for test samples on the inertial stand is presented in Fig. 9. The values obtained by simulating the conditions of real friction, on the inertial test-stand, were brought close
Fig. 3. Sintered friction materials (second lot), fixed on a metallic pad, for test samples on inertial stand.
In this design, n = 2 (i.e. applied load and relative speed in braking point) and P = 2 (i.e. upper and lower level of each variable). Thus, the minimum number of trial experiments to be conducted for each material is four. If wear rate is represented by Y1 and Y2 (Y1 measured by height loss of material and Y2 measured by weight loss) and friction coefficient is represented by Y3 , the linear regression equation for these experiments could be written as: Yi = bi0 + bi1x1 + bi2x2 + bi3x1x2
(5)
where bi0 is the response variable of wear rate and friction coefficient, respectively, at the base level (i.e. at the applied load 7.45 × 10−1 MPa and relative speed in braking points 14.7 m/s); bi1 and bi2 are coefficients associated with each variable x1 (applied load) and x2 (relative speed in braking point) and bi3 is interaction coefficient between x1 and x2 , within the selected levels of each of the variables. The parameters of Eq. (5) have been estimated by the method of least squares using a Excel programme package. The coefficients were determined statistically using the Student criterion and the accordance degree of the regression equations was verified with the Fischer criterion. The mathematical model of wear behaviour once established, the problem of the models optimization is to determination of the mathematical function extreme [30].
Table 2 The physical characteristics of raw powders. Powder type
Fe Cu Graphite MoS2 Mullite (3Al2 O3 .2SiO2 ) Spinel (MgAl2 O4 )
Particle size distribution, % >100 m
100–63 m
63–50 m
50–40 m
<40 m
23.4 – – – – –
28.1 54.4 69.8 44 – –
28.3 24.7 20.4 29 23 36
11.9 12.2 5.1 12.5 42 32.1
8.3 8.7 4.7 14.5 35 31.9
Apparent density, g/cm3
Flow rate, s/50 g
2.55 1.34 0.49 0.43 0.72 0.81
39 41.2 Not flowing Not flowing Not flowing Not flowing
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Fig. 4. Morphology of (a) Fe powder (1000×), (b) Cu powder (1000×), (c) graphite powder (3000×), (d) MoS2 powder (2000×) and (e) Mullite powder (1000×).
to the maximal allowed amount of 0.5–0.6 at the cold testing, under speed increasing conditions (Fig. 9(b)). 4.3. Linear regression equations The upper level and lower level of each variable along with their code values used in these investigations are presented in Table 4. Experiment planning matrix, the values of individual variables with their wear rate and friction coefficient response in each trial tested for SFM1, SFM2 and RM samples (results of random experiments along with theoretical value) are presented in Tables 5–7 (where y1 , y2 and y3 are the experimental values and y1calc , y2calc and y3calc are the calculated ones). The calculation of dispersal reproducibility for y1 , y2 – wear rate and y3 – friction coefficient was made in cases in which the values of Z1 xZ2 have the base level.
The four parallel experiences are presented in Table 8, where u is the number of experience; n is the number of parallel experiments is the arithmetic average of obtained results for k = 1, 2 or (n = 4); yuk 3, the type of variable response; v2 is the degrees of freedom at n − 1. The calculation of reproducibility dispersion is based on the Eq. (6):
n S02
=
u=1
yu
n−1
(6)
Result for variable response Y1 is S02 = 0.00071875, for variable response Y2 is S02 = 0.000583 and for variable response Y3 is S02 = 0.000475, for sample SFM1. The coefficients were determined statistically using the Student criterion and the accordance degree of the regression equations was verified with the Fischer criterion.
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Fig. 5. The structure of iron–copper based composites (SFM1): (a) microstructure ESEM, (b) EDS elemental analysis, and (c) compo image of samples, 500× magnification.
Thus, the coefficients of each variable from Eq. (5) were calculated with Eqs. (7) and (8).
n
x · yu u=1 iu , (xiu )2
bi =
n bij =
x u=1 iu
for linear effects,
· xju · yu
(xiu · xju )2
,
for interaction effects,
(7)
(8)
By introducing the calculated coefficients (7) and (8) in Eq. (5) the following regression equations result: Y1,SFM1 = 1.37 + 0.45x1 + 0.07x2 + 0.03x1 x2
(9)
Y2,SFM1 = 0.535 + 0.101x1 + 0.037x2 − 0.0325x1 x2
(10)
Y3,SFM1 = 0.454 − 0.083x1 − 0.0138x1 x2
(11)
where Y1,SFM1 , Y2,SFM1 represent the response variables (wear rate of SFM1 samples measured by height loss and weight loss, respectively; Y3,SFM1 represents the friction coefficients of SFM1). The dispersion in determination of coefficients is calculated with relations: 2 Sbi =
S02
x j=1 ij
and Sbi =
2 Sbi
(12)
Table 4 Levels of each factor and their coded values. SI. no.
1 2 3
Variables Factor levels
Z1 – Applied load (×10−1 MPa)
Z2 – Relative speed in braking point (m/s)
Coded values Upper level (+1) Base level, Z10 (0) Lower level (−1) Variation interval, Z1
x1 11.5 7.5 3.5 4
x2 16.9 14.7 12.5 2.2
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Fig. 6. The X-ray mapping of the elements (a) and EDS elemental analysis of the frictional sintered iron–copper based material (SFM1) (b).
Result for variable response Y1 , Sbi = 0.013404757, for variable response Y2 , Sbi = 0.012076, and variable response Y3 , Sbi = 0.010897. The Student criterion t0.05;4 = 2.776 and bi – confidence interval is bi = 2.776·Sbi . After calculation, result for variable response Y1,SFM1 , bi = 0.0372, for variable response Y2,SFM1 , bi = 0.0335 and for variable response Y3,SFM1 , bi = 0.0302. But according to [30], if the relation b0 , b1 bi is not accomplished, result for the calculated value of the coefficient is not statistically different
from zero; therefore, the term comprising interaction, or another factor from equation, will be neglected. In this case, the calculated (predicted) linear model of the processes will be: for the wear, Eqs. (13) and (14) and friction coefficients Eq. (15), respectively, of the friction material SFM1 can be expressed as follows: Y1,SFM1 = 1.37 + 0.45x1 + 0.07x2
(13)
Y2,SFM1 = 0.535 + 0.101x1 + 0.037x2
(14)
Y3,SFM1 = 0.454 − 0.083x1
(15)
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Fig. 7. The microstructure ESEM (SFM1), 500× maginification (a) and EDS elemental analysis taken from the compact white phase (b), from small white particles (c) and light grey compact particles (d).
In this equation x1 and x2 are coded values of the factors, which are calculated from natural values, with the formula: xi =
zi − zio zi
(16)
Substituting in Eq. (16), the values from Table 4, we will z −7.5 z −14.7 obtain(17)x1 = 1 4 and x1 = 2 2.2
Verification of the model concordance: The computation of dis2 persion produced by the linear equations of regression Sconc is presented in Table 9, where
yu = yexp − y
Fig. 8. The microstructure ESEM (SFM1), 2000× maginification (a) and EDS elemental analysis taken from the light grey compact particles (b).
(18)
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Fig. 9. The variation friction coefficient depending on temperature, speed and number of brakes for test samples on the inertial stand in: grinding (a), cold testing (b), hot testing (c), recovery testing (d), and endurance conditions (e).
and n 2 Sconc =
(yu )
2
u=1
1
(19)
The calculated Fischer criterion is: Fc =
2 Sconc
S02
(20)
Introducing the values of yu from Table 9 in Eq. (19) and then 2 introducing the resulted Sconc and values of reproducibility disper2 sion S0 in Eq. (20) result that the calculated Fisher is Fc1 = 5.01 (for variable response Y1 ); Fc2 = 7.307 (for variable response Y2 ) and Fc3 = 5.05 (for variable response Y3 ). From tabulated values of optimization book annexes in [29] resulted in F˛ , 1 , 2 = F0.05,1.3 = 10.13. Because Fci < F0.05,1.3 , i = 1.2 or 3 ⇒. The linear models for all three variables of sintered friction material SFM1 are in concordance with analyzed process.
Table 5 The values of individual variables with their wear rate and friction coefficient response in each trial tested for SFM1 samples. Trial no.
1 2 3 4
x0
Applied load (×10−1 MPa) x1
Relative speed (m/s) x2
(+1) (+1) (+1) (+1)
11.5 (+1) 3.5 (−1) 11.5 (+1) 3.5 (−1)
16.9 (+1) 16.9 (+1) 12.5 (−1) 12.5 (−1)
x1 x2
Wear rate (mm/h) y1
Calculated wear rate (mm/h) y1calc
Wear rate (g/cm2 h) y2
Calculated wear rate (g/cm2 h) y2calc
Friction coefficient y3
Calculated friction coefficient y3calc
(+1) (−1) (−1) (+1)
1.92 0.96 1.72 0.88
1.89 0.99 1.75 0.85
0.64 0.503 0.631 0.364
0.673 0.471 0.599 0.397
0.337 0.53 0.405 0.543
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Table 6 The values of individual variables with their wear and friction coefficient response in each trial tested for SFM2 samples. Trial no.
1 2 3 4
x0
Applied load (×10−1 MPa) x1
Relative speed (m/s) x2
(+1) (+1) (+1) (+1)
11.5 (+1) 3.5 (−1) 11.5 (+1) 3.5 (−1)
16.9 (+1) 16.9 (+1) 12.5 (−1) 12.5 (−1)
x1 x2
Wear rate (mm/h) y1
Calculated wear rate (mm/h) y1calc
Wear rate (g/cm2 h) y2
Calculated wear rate (g/cm2 h) y2calc
Friction coefficient y3
Calculated friction coefficient y3calc
(+1) (−1) (−1) (+1)
1.4 1.1 1.25 0.86
1.424 1.078 1.228 0.882
0.52 0.29 0.368 0.26
0.491 0.321 0.399 0.229
0.295 0.395 0.32 0.46
0.295 0.395 0.32 0.46
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0.353 0.529 0.287 0.463 0.39 0.49 0.25 0.5 0.723 0.317 0.759 0.353 0.65 0.39 0.83 0.28 1.559 1.333 1.433 1.207 1.56 1.33 1.43 1.21 (+1) (−1) (−1) (+1) 16.9 (+1) 16.9 (+1) 12.5 (−1) 12.5 (−1) 11.5 (+1) 3.5 (−1) 11.5 (+1) 3.5 (−1) (+1) (+1) (+1) (+1)
x1 x2 x0
In the same way the 4 parallel experiences for variable response for SFM2 and RM samples were calculated and the following linear regression equations (21)–(26) were obtained which are also in concordance with analyzed process:
1 2 3 4
Calculated wear rate (mm/h) y1calc Wear rate (mm/h) y1 Relative speed (m/s) x2 Applied load (×10−1 MPa) x1 Trial no.
Table 7 The values of individual variables with their wear and friction coefficient response in each trial tested for RM samples.
Wear rate (g/cm2 h) y2
Calculated wear rate (g/cm2 h) y2calc
Friction coefficient y3
Calculated friction coefficient y3calc
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Y1,SFM2 = 1.153 + 0.173x1 + 0.098x2
(21)
Y2,SFM2 = 0.359 + 0.085x1 + 0.046x2
(22)
Y3,SFM2 = 0.368 − 0.06x1
(23)
Respectively: Y1,MR = 1.383 + 0.113x1 + 0.063x2
(24)
Y2,MR = 0.538 + 0.203x1 − 0.073x1 x2
(25)
Y3,MR = 0.408 − 0.088x1
(26)
From Eq. (21) deduced that the height loss increase with increasing of applied load, relative speed in braking point and no depends with interaction of those variables. Eq. (22) showed that the weight loss of SFM2 material depends direct proportional with applied load and no depend of relative speed (witch means that regardless of automotive speed, braking to accomplish in safe condition depending on pressing force. Eq. (23) showed that the friction coefficient of SFM2 material increased with increased all the variables inclusive with interaction of them. We noticed that wear rate measured by height loss for reference material (Y1,MR ), weight loss of material increased with increasing of applied load and relative speed in accordance with Eq. (24) and wear rate measured by weight loss (Y2,MR ) increased with increasing of applied load and with decreasing of interaction of applied load and relative speed, in accordance with Eq. (25). Eq. (26) shows that the friction coefficient of reference material decreases with increasing of applied load. The calculated (theoretical) values of wear rate and friction coefficient were obtained replacing in Eqs. (13)–(15) and (21)–(26) the coded values corresponding of each trial for all types of material tested. Through the statistical experimental design presented before, we can describe the evolution of tribological characteristics of sintered frictions samples (SFM1, SFM2 and RM) at different applied load, different relative speed and constant temperature. Thus, for comparison, the statistical results of response variables, represented by wear rates and friction coefficients, and the experimental data of them, are presented in Figs. 10 and 11. For better understanding of tribological behaviour of the three types of samples depending on all factors: applied load (3.5 × 10−1 MPa and 11.5 × 10−1 MPa) and relative speed (16.9 m/s and 12.5 m/s), we presented all the results of experimental and calculated data in Fig. 11. From these figures we observed that the difference between experimental values and the calculated one are the majority ±3–6%. Comparing the wear rates of SFM1 and SFM2 samples with reference material RM we noticed that values for all materials are not great; in addition at maximum values of applied load and relative speed, wear rate measured by loosing weight per hour is 0.64 g/cm2 h for SFM1, 0.32 g/cm2 h for SFM2, 0.65 g/cm2 h for MR, and wear rate measured by loosing height per hour is 1.92 mm/h for SFM1, 1.4 mm/h for SFM2 and 1.56 mm/h for RM. We have seen that the lowest values of wear rate have friction materials SFM2, which have a lowest value in comparison with the reference material. The values of friction coefficients at maximum solicitations are:
I.N. Popescu et al. / Applied Surface Science 285P (2013) 72–85 Table 8 The 4 parallel experiences for variable response for SFM1 sample.
No.
y ’ u1
yu1 = y u1 − yu1
(yu1 )
1 2 3 4
1.52 1.6 1.64 1.43
0.0275 0.0525 0.0925 0.1175
0.00075625 0.00275625 0.00855625 0.01380625
2
y ’ u2
yu2 = y u2 − yu2
(yu2 )
0.51 0.48 0.59 0.5
0.01 0.04 0.07 0.02
0.0001 0.0016 0.0049 0.0004
2
83
y’3
yu3 = y u3 − yu3
(yu3 )
0.47 0.42 0.47 0.38
0.035 0.015 0.035 0.055
0.001225 0.000225 0.001225 0.003025
2
2 3 3 3 3
Table 9 2 produced by the linear equation of regression, for SFM1 sample. The computation of dispersion Sconc No.
y1exp
y¯ 1
y1u
(y1u )2
y2exp
y¯ 2
y2u
(y2u )2
y3exp
y¯ 3
y3u
(y3u )2
1 = N − k
1 2 3 4
1.92 0.96 1.72 0.88
1.89 0.99 1.75 0.85
0.03 0.03 0.03 0.03
0.0009 0.0009 0.0009 0.0009
0.64 0.503 0.631 0.364
0.673 0.471 0.599 0.397
0.033 0.032 0.032 0.033
0.001089 0.001024 0.001024 0.001089
0.337 0.53 0.405 0.543
0.371 0.537 0.371 0.537
0.034 0.007 0.034 0.006
0.001156 0.000049 0.001156 0.000036
1 1 1 1
0.337 for SFM1, 0.36 for SFM2, 0.39 for MR respectively. We noticed that SFM1 has a constant and higher value of friction coefficient; this value is closer to a coefficient of friction value of MR material. From Figs. 10 and 11, we can see that the values of friction coefficients have a low decrease with increasing of strains (load and speed), decreasing from 0.543 to 0.337 for SFM1, from 0.46 to 0.295for SFM2, from 0.5 to 0.39 for RM (maximum value was obtained for applied load of 3.5 × 10−1 MPa and 16.9 m/s speed);
all values of frictional coefficients are in admitted domain (0.5–0.3) for automobile friction materials. From statically point of view, after we analyzed Eqs. (13)–(15) for SFM1 sintered friction material, Eqs. (21)–(23) for SFM2 sintered friction material and Eqs. (24)–(26) for reference material RM respectively were found them “in accordance” on the basis of the Fischer criterion which means that the model is good and supports the reproducibility of the trial experiments.
Fig. 10. The experimental and statistical calculated values of response variables: (a) wear rates of samples as a function of applied load, at constant speed (16.9 m/s), (b) friction coefficients of samples as a function of applied load, at constant speed (16.9 m/s), (c) wear rates of samples as a function of applied load, at constant speed (12.5 m/s); (d) friction coefficients of samples as a function of applied load, at constant speed (12.5 m/s).
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Fig. 11. The experimental and calculated values of wear rate and friction coefficient respectively of all three types of materials as a function of applied load (3.5, 11.5 × 10−1 MPa, respectively), for relative speed (a) 16.9 m/s and (b) 12.5 m/s.
5. Conclusions We obtained a sintered friction iron–copper based material, with new chemical compositions and proportion of components (SFM1 and SFM2 samples) and non-polluting constituents (ecologically friendly). The chosen composition and proportion of these components ensured good physical and mechanical characteristics, confirmed also by the morphological and microcompositional investigation (uniform distribution of constituents and good adhesion between them, due to changes occurring during liquid phase sintering in the Fe–Cu system). We established after different conditions of tribological tests, including the tests on inertial stand (inertia dynamometer) which simulate the real conditions of dry friction, that the obtained frictional materials for braking pad have: (i) high and stable friction coefficients (0.3–0.5), according to the requirements imposed by brake pads materials for automotive application (Fig. 9); (ii) high durability of SFM2 materials in comparison with RM and SFM1 samples; the values of wear rate, in the same experimental conditions for SFM2 is maximum 25% lower than RM (Figs. 10 and 11). (iii) increase of applied load or increase of the relative speed in braking points induces moderate increase of wear rate and slow decrease of the frictional coefficients of the composites (shown through experimental data and obtained linear regression equations in Figs. 9–11). Acknowledgements The authors are grateful to the Metallurgical Research Institute Bucharest and BIOMAT Research Centre, Bucharest for providing equipments and also thank Full Professor Ph.D. Dionezie Bojin from Electron Microscopy Laboratory, Science and Engineering Faculty, POLITEHNICA University of Bucharest for advice and support. References [1] H. Jang, K. Ko, S.J. Kim, R.H. Basch, J.W. Fash, The effect of metal fibers on the friction performance of automotive brake friction materials, Wear 256 (3–4) (2003) 406–414.
[2] H. Zaidi, A. Senouci, Thermal tribological behaviour of composite carbon metal/steel brake, Applied Surface Science 144–145 (1999) 265–271. [3] N. Popescu, E. Geza, Contribution to increase of the performant composite materials having friction properties, in: Proceedings of the 2nd National Powder Metallurgy Conference, Metallurgical and Materials Engineering Dept., Ankara, Turkey, 15–17 July, 1999, pp. 225–233. [4] P.J. Blau, H.M. Meyer III, Characteristics of wear particles produced during friction tests of conventional and unconventional disc brake materials, Wear 255 (2003) 1261–1269. [5] M. Eriksson, F. Bergman, S. Jacobson, Surface characterization of brake pads after running under silent and squealing conditions, Wear 232 (1999) 163–167. [6] Y.M. Wang, B.L. Jiang, L.X. Guo, T.Q. Lei, Tribological behavior of microarc oxidation coatings formed on titanium alloys against steel in dry and solid lubrication sliding, Applied Surface Science 252 (2006) 2989–2998. [7] A.E. Anderson, Friction and wear of automotive brakes, in: ASM handbook, vol. 7. Powder Metallurgy, ASM International, USA, 1984, pp. 569–577. [8] D. Chan, G.W. Stachowiak, Review of automotive brake friction materials, Proceedings of the Institution of Mechanical Engineers, Part D: Journal of Automobile Engineering 218 (2004) 953–966. [9] M.L. Marucci, HanejkoF F.G., Effect of Copper Alloy Addition Method on the Dimensional Response of Sintered Fe–Cu–C Steels, 2010, pp. 1–11 http://www.mpif.org [10] T. Murphy, M. Baran, An investigation into the effect of copper and graphite additions to sinter hardening steels, in: Advances in Powder Metallurgy and Particulate Materials, part 10, Metal Powder Industries Federation, Princeton, NJ, 2004, pp. 266–274. [11] M. Morakotjinda, R. Krataitong, P. Wila, P. Siriphol, A. Daraphan, O. Coo, Sintered Fe-Cu-C materials, Chiang Mai Journal of Science 35 (2) (2008) 258–265. [12] V. Prasad, P.K. Rohatagi, Tribological properties of al alloy particle composite, Journal of Metals 11 (1987) 22. [13] X. Xiong, H-C. Sheng, C. Jie, Y. Ping-ping, Effects of sintering pressure and temperature on microstructure and tribological characteristic of cu-based aircraft brake material, Transactions of Nonferrous Metals Society of China 17 (2007) 669–675. [14] E. Hounkponou, H. Nery, D. Paulmier, A. Bouchoucha, H. Zaidi, Tribological behaviour of graphite/graphite and graphite/copper couples in sliding electrical contact: influence on the contact electric field on the surface passivation, Applied Surface Science 70–71 (1) (1993) 176–217. [15] Popescu IN, Ernszt G. Researches Projects Orizont 2000 No. C 55B, Theme A33/3. Composites materials with iron based matrix used for frictional materials (in Romanian). [16] I.N. Popescu, S. Zamfir, V.F. Anghelina, C.O. Rusanescu, Processing by P/M route and characterization of new ecological aluminum matrix composites (AMC), International Journal of Mechanics 3 (4) (2010) 71–80. [17] I.N. Popescu, S. Zamfir, D. Bojin, M. Brânzei, D. Gheorghe, L.A. Sorcoi, Physical and quantitative microstructural analysis of sintered Al–Cu/SiCp composites, Materials Science Forum 672 (2011) 251–254. [18] S. Chen, C. Kang, J. Wang, C. Liu, K. Sun, Synthesis of anodizing composite films containing superfine Al2 O3 and PTFE particles on Al alloys, Applied Surface Science 256 (2010) 6518–6525. [19] V. Tsakiris, W. Kappel, E. Enescu, G. Alecu, F. Albu, F. Grigore, V. Marinescu, M. Lungu, Characterization of Al matrix composites reinforced with
I.N. Popescu et al. / Applied Surface Science 285P (2013) 72–85
[20]
[21]
[22]
[23]
[24]
alumina nanoparticles obtained by pm method, Journal of Optoelectronics and Advanced Materials 13 (9) (2011) 1172. M. Lucaci, M. Valeanu, R.L. Orban, V. Tsakiris, D.C. Cristea, L. Leonat, Shape memory NITI alloys obtained by powder metallurgy route, Materials Science Forum 672 (2011) 99–104. V. Andrei, E. Andrei, G.H. Vlaicu, C. Stihi, G. Dima, C. Oros, S. Dinu, Chemical binding and structure of carbonic thin films with advanced properties studied by electron spectroscopy, Journal of Optoelectronics and Advanced Materials 9 (7) (2007) 2291–2295. I.N. Popescu, D. Bojin, I. Carceanu, G. Novac, F.V. Anghelina, Morphological and structural aspects using electronic microscopy and image analysis of iron powders obtained by water atomization process, Journal of Science and Arts 1 (12) (2010) 125–134. S. Domsa, M. Firanescu, Influence of sintering conditions on properties of sintered friction materials based on iron powder, Metallurgy and New Materials Researches 18 (1997) 655. R.I. Zamfir, The influence of the laser surface melting parameters (LSM) on the structure and microhardness of the 316L austenitic stainless steel, UPB Scientific Bulletin 71 (2009) 137.
85
[25] J. Hernáez, E. Otero, A. Pardo, C. Merino, M. Laguna, Contribution to the microstructural study of Fe-8% Cu sintered compacts, Revista de Metalurgia (Madrid) 22 (2) (1986) 95–101. [26] D. Bojin, D. Bunea, Fl. Minculescu, M. Minculescu, Scanning Electron Microscopy and Applications, AGIR, Bucharest, 2005 (in Romanian). [27] Y. Sahin, Wear behavior of aluminium alloy and its composites reinforced by SiC particles using statistical analysis, Materials & Design 24 (2003) 95–103. [28] I.N. Popescu, V. Bratu, M.C. Enescu, Experimental researches and statistical analysis of the corrosion behavior of rolled and heat treated 2XXX Al alloys, in: AEE’10 Proceedings of the 9th WSEAS, International Conference on Applications of Electrical Engineering, Recent Advances in Electrical Engineering (2010) 225–232. [29] C. Ghit¸a˘ , Mathematical Metallurgy, Romanian Academy, Bucharest, 1995 (in Romanian). [30] D. Taloi, E. Florian, C. Bratu, E. Berceanu (Eds.), The Optimisation of Metallurgical Processes, Didactica & Pedagogic˘a, Bucharest, 1983 (in Romanian).
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Applied Surface Science journal homepage: www.elsevier.com/locate/apsusc
Tribological behaviour and statistical experimental design of sintered iron–copper based composites Ileana Nicoleta Popescu a , Constantin Ghit¸a˘ b , Vasile Bratu a,∗ , Guillermo Palacios Navarro c a b c
Valahia University of Targoviste, Faculty of Materials and Mechanics Engineering, 18-24 Unirii Boulevard, 130082, Targoviste, Romania Valahia University of Targoviste, Faculty of Science and Arts, 18-24 Unirii Boulevard, 130082, Targoviste, Romania University of Zaragoza, Department of Electronic Engineering and Communications, Zaragoza, Spain
a r t i c l e
i n f o
Article history: Received 12 April 2013 Received in revised form 30 July 2013 Accepted 1 August 2013 Available online 11 August 2013 Keywords: Tribological characteristics ESEM EDS X-ray maps and compo image of surface friction materials Statistical experimental design
a b s t r a c t The sintered iron–copper based composites for automotive brake pads have a complex composite composition and should have good physical, mechanical and tribological characteristics. In this paper, we obtained frictional composites by Powder Metallurgy (P/M) technique and we have characterized them by microstructural and tribological point of view. The morphology of raw powders was determined by SEM and the surfaces of obtained sintered friction materials were analyzed by ESEM, EDS elemental and compo-images analyses. One lot of samples were tested on a “pin-on-disc” type wear machine under dry sliding conditions, at applied load between 3.5 and 11.5 × 10−1 MPa and 12.5 and 16.9 m/s relative speed in braking point at constant temperature. The other lot of samples were tested on an inertial test stand according to a methodology simulating the real conditions of dry friction, at a contact pressure of 2.5–3 MPa, at 300–1200 rpm. The most important characteristics required for sintered friction materials are high and stable friction coefficient during breaking and also, for high durability in service, must have: low wear, high corrosion resistance, high thermal conductivity, mechanical resistance and thermal stability at elevated temperature. Because of the tribological characteristics importance (wear rate and friction coefficient) of sintered iron–copper based composites, we predicted the tribological behaviour through statistical analysis. For the first lot of samples, the response variables Yi (represented by the wear rate and friction coefficient) have been correlated with x1 and x2 (the code value of applied load and relative speed in braking points, respectively) using a linear factorial design approach. We obtained brake friction materials with improved wear resistance characteristics and high and stable friction coefficients. It has been shown, through experimental data and obtained linear regression equations, that the sintered composites wear rate increases with increasing applied load and relative speed, but in the same conditions, the frictional coefficients slowly decrease. © 2013 Elsevier B.V. All rights reserved.
1. Introduction The increasing consumer demand for low-cost performing and non-polluting materials, with a favourable price/quality ratio, has determined, in the past decades, development of extensive researches in automotive brake friction industry. The efforts include the development of ferrous (composites with iron based matrix) and non-ferrous based materials such as copper, aluminium or carbon composites, as new candidates of brake friction materials [1–6]. Friction materials are used for converting the kinetic energy of the automotive to thermal energy (heat), resulting from two surfaces in contact of the brake pads and brake disc, which is absorbed and dissipated by the materials [7].
∗ Corresponding author. Tel.: +40 245206106. E-mail address: vasilebratu.uvt [email protected] (V. Bratu). 0169-4332/$ – see front matter © 2013 Elsevier B.V. All rights reserved. http://dx.doi.org/10.1016/j.apsusc.2013.08.007
The brake friction materials have a complex composition that includes both metallic and non-metallic components. Also, brake pads typically comprise the binders, the lubricating components and frictional additives [7]. The purpose of a binder is to maintain the brake pads’ structural integrity (to hold the components of a brake pad together) under mechanical and thermal stresses. The binders could be synthetic resins, metallic ingredients or mixture of metallic and organic ingredients. The synthetic resin based friction materials is cheap to produce, but it has as a disadvantage that, in high-energy braking applications (operating up to 370–450 ◦ C) when the induced temperature can be high enough, the material decomposes or breaks down. As a result, the friction coefficient with the brake disc is compromised leading to decreasing road safety [7,8]. In case of using metallic binders as iron, cooper, aluminium and their alloys, we can improve the wear resistance, thermal diffusivity and strength of the brake friction material. The lubricating components such as graphite, molybdenum disulphide, etc. have the objective to: (i) stabilize the developed friction coefficient
I.N. Popescu et al. / Applied Surface Science 285P (2013) 72–85
during braking, particularly at high temperatures, (ii) decrease wear of counterparts and also (iii) increase grippe resistance [1–14]. The frictional additives, such as oxides (Al2 O3 , SiO2 , ZrO2 , Mullite, and Spinel), nitrides (TiN and Si3 N4 ), carbides, (SiC, TiC, B4 C, and VC), determine the frictional properties of the brake pads. Also the frictional components increase and stabilize the friction coefficient values, respectively increase the wear resistance [1–12,15]. Composites with iron or steel matrix and ceramic reinforcements used like friction materials bring new possibilities in the production of wear resistant materials because they provide good wear resistance and maintain friction effectiveness at elevated temperature [7,11,15]. Iron could be alloyed with Cu, Cu–Sn–(Zn), Ni, Mo, Al, Co, Mn, etc. for improving resistance of the metallic matrix [7–11,15]. Cooper and Cu alloys are chosen mainly to improve the thermal conductibility at the friction interface and for sustaining the level of the coefficient of friction at elevated temperatures by producing copper oxides at the friction interface [1,8]. Aluminium or aluminium alloys are used in frictional materials because of lightweight and good thermal diffusivity [12,16–18]. The proportion and granulation of these components, was and is still the subject of numerous investigations in various countries in order to achieve the optimal characteristics of brake friction materials. The processing of brake friction materials is realized by a diversity of methods [6–8,13–23] including powder metallurgy (P/M) techniques. The P/M techniques give (i) a good uniformity distribution of lubricated and frictional components in metallic matrix, (ii) near net-shape parts fabrication, (iii) high productivity (iv) dimensional accuracy and (v) allows a good control of brake friction require characteristics. Among the required market performance criteria for brake pads materials, we mention [8]: (a) maintaining a high and stable friction coefficient with the brake disc, in various conditions, including at high temperatures; (b) very good wear resistant material and implicit very good durability in service; (c) good strength at elevated temperatures; (d) high thermal conductivity; (e) high corrosion resistance; (f) smooth braking assurance. In this paper we obtained frictional composites by conventional P/M technique and we have characterized them by microstructural and tribological point of view.
2. Materials and compositional selection The raw powders chosen for frictional materials were Fe (reduced), Cu (electrolytic) as metallic ingredients, graphite and MoS2 as solid lubricant, respectively, mullite and spinel as frictional components, with a particles size situated up to 100 m. The raw powders selected must be non-polluting material from compositional point of view and have required characteristics. Using iron/steel based material in comparison with Cu-based material improved strength, hardness, ductility and heat resistance properties. Iron as a matrix of the automotive brake material has been chosen due of its stability under higher temperatures (1100 ◦ C), which imply very good tribological behaviour under heavy duty dry operating conditions. The selection of copper as alloying element gives additional strength of sintered steels (the elemental copper at 1083 ◦ C promotes sintering and enhances the strength of the steel) [10]. In order to improve cold and hot strength of sintered friction materials a tight ratio between cooper and iron was chosen. Generally for the friction sintered materials for automotive brakes, the cooper ratio in the material is majority [3]. Graphite + MoS2 have the role to stabilize the developed friction coefficient during braking, particularly at elevated temperatures. Additions of graphite to Fe–Cu sintered steels are desirable because the carbon promotes the formation of a pearlitic microstructure, resulting in additional strength and hardness of steel [9].
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The high proportion in lubricating components was chosen due to (i) easy pressing of the mixture, at medium pressure of 500–600 MPa; (ii) improvement of sliding properties, by forming a lubricant layer on the opposing counter friction material at speed of 20–30 m/s [3,7,14]. The new chemical compositions (new proportion of components) established after a number of preliminary tests [15] of iron–copper based friction material (SFM1 and SFM2) samples preparation and the chemical composition of a commercial brake pad, used as reference material (RM) are presented in Table 1. 3. Materials and experimental procedure 3.1. Powder analysis The following characteristics of powders were analyzed: (i) the apparent density according to SR EN 23923-1: 1998 standards, (ii) flow rate, in concordance with the standard method SR EN ISO 3953: 1998; (iii) particle size distribution, according with SR 1320394 and ISO 4497:1983; (iv) the morphology of powder particles determined by scanning electronic microscopy (SEM). 3.2. Composite preparation We used P/M as a technique for preparation of composites for automotive brake pads because this technology allows us to obtain sintered friction materials with a large variety of chemistry in concordance with different conditions of solicitation. Before the manufacturing of the composite, the oxides of iron copper powders were reduced in “Siemens–Plania” type furnace in presence of H2 atmosphere at temperature near 980 ◦ C for iron and 280 ◦ C for copper, during 60 min, in order to obtain reduced metallic powders. The ceramic powders were heated for 300–400 ◦ C, holding 120 min, for elimination of adsorbed gases, moisture and organic contaminants. The dosage of mixtures was made gravimetric and the elemental powder of the mixtures was dry blended using the Double Cone Blender (10 kg capacity) at a rotation speed of 20 rpm, during 6–8 h. The mixed homogenous powders were compacted at room temperature in a double action hardened steel die with a automate hydraulic press of 30 tone force, Meyer Type, at pressure of 500–600 MPa. According to literature [23–25] there are a diversity of process parameters (sintering temperature, sintering time, and protective environment). For our experiments, the compacts were sintered in a sintering furnace, Balzers type, at 1050 ◦ C sintering temperature, 90 min holding time, in vacuum atmosphere under a pressure of 0.1–0.3 mmHg, in limits mentioned of scientific literature. The composite processing flow-sheet is presented schematically in Fig. 1. 3.3. Sintered friction material characterization The obtained sintered SFM1, SFM2, composites and reference material RM were analyzed from physical, mechanical, microstructural (SEM) and micro-compositional (SEM, EDS and X-ray maps) point of view, according with specific techniques presented in [26]. The quantitative X-ray microanalysis of obtained sintered iron–copper based composite as frictional material was carried out using the energy dispersive X-ray spectroscopy (EDS) of the environmental scanning electron microscopy (ESEM), FEI XL-30 and FEG SEM Inspect F, FEI. The density of the sintered composites was measured by physical measurements and also using the Archimedes method. The Brinell Hardness test was done according ISO 4498/1 standardization.
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Table 1 Raw powders chemical composition (wt.%) of brake friction materials. Friction material type
SFM1 SFM2 RM
Metallic binder
Lubricant component
Frictional components
Fe
Cu
Sn
Graphite
MoS2
Mullite (3Al2 O3 ·2SiO2 )
24–30 30–35 30–32.5
21–25 20–30 30–32.5
– – 2–3
18–22 13–17 9–11
4–8 5–7 4–6
The tribological (friction and wear) tests of sintered iron–copper based composite for automotive brake pads were performed using a “pin-on-disc” tribometers. One lot was tested with a tribometer (with a grey cast iron of 300 HB as counterparts), on 10 mm × 10 mm × 7 mm samples, in dry friction conditions, applied normal load between 0.35 and 1.15 MPa, during 15–20 min per sample, at 12.5–16.9 m/s relative speed in braking points (sliding distance between 8125 and 14733 m) and at constant temperature. The test machine consists of a balanced pendulum in which is mounted the sample into a box. The sample rubs on a cast iron disc and the disc is acted through the main body of a transmission belt, by an electric motor of 1.4 kW power. The peripheral speed of the disc can be varied by positioning the sample box (sample position) at different distances from the axis of the machine through a variable ratio transmission. The frictional force was determined by measuring the angular movement (˛) of pendulum from the vertical position (Fig. 2). The correlation between applied pendulum weight (G) in Kg, total length (L) in mm and the distance from the spindle axis to pendulum test box (e) in mm is given by the relation. F = L · G · sin ˛/e
(1)
Spinel (MgAl2 O4 )
Balanced Balanced Balanced
−
From Eqs. (1) and (2) result that the coefficient of friction is determined by the relationship: = L · G · sin ˛/Ne
(3)
The wear rate was measured, respectively, by height loss (mm/h) and weight loss (g/cm2 h). The wear of the friction materials was expressed by weight/height loss percent as: W % = (W0 − W1 )/W0 · 100%
(4)
where W0 and W1 are the weight/height of the specimens before and after testing, respectively. The other lot of samples were tested on an inertial test stand according with Fermit S.A. methodology [15]. In this type of testing the real conditions of dry friction of automotive brake pads were simulated (cold and hot testing, grinding and endurance conditions) on dry sliding wear conditions, at a contact pressure of 2.5–3 MPa, depending on braking parameters as: (i) number of braking stops (25–500) and (ii) speed 300–1200 rpm. The temperature during braking ranged between 75 and 350 ◦ C. The dimensions of sintered friction samples of the second lot, for test samples on inertial stand, fixed on a metallic pad, were 2500 mm2 × 12 mm (Fig. 3).
It is known that: = F/N
(2)
where F is the frictional forces; is the coefficient of friction and N is the normal forces.
Fig. 1. The P/M technology route of SFM1 and SFM2 materials for braking pads.
3.4. Factorial design Because of the tribological characteristics importance (wear rate and friction coefficient) of sintered iron–copper based composites, we predicted the tribological behaviour through statistical analysis [27–30]. For that purpose, a factorial design of experiment of the type Pn [30] was used in the present paper, for the testing sintered frictional material from the first lot, in dry friction conditions. The “n” corresponds to the number of factors and “P” stands for the number of levels.
Fig. 2. Schematic diagram of the forces which activate during the wearing of samples (pin-on-disc).
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Table 3 The average values concerning physical and mechanical characteristics of iron–copper sintered composites for automotive brakes. Sintered friction material
Density, g/cm3
Hardness, HB
SFM1 SFM2 RM
3.8 3.9 3.6
16.5 19.2 11.8
4. Results and discussion
graphite and MoS2 have the particles shape as flaws, especially the molybdenum disulphide particles have the tendency of coalescence. The average values of apparent densities of the mixtures for frictional materials were 1.39 g/cm3 for SFM1 powder mixture and 1.46 g/cm3 for SFM2 mixture. The physical and mechanical characteristics (Brinell hardness and density) of the obtained composite friction material, in comparison with reference material, are presented in Table 3. The structure of iron–cooper composites, the ESEM, EDS analysis and X-ray mapping of the elements from SFM1 samples are presented in Figs. 5–8. The analyses were made on the surface of the frictional composite material, from different white and grey compact areas. In Figs. 5 and 6 we can observe the components and elements from the surface of sintered iron–copper based composite. The microstructure and elemental analysis shows a basic mass formed by iron and copper (white compact areas) as we observed in Fig. 5(a) and (c), respectively and in Fig. 7(a) and (b). The mechanical inclusions, especially mullite, spinel, are identified from X-ray mapping (Al, O, Si, and Mg elements) and EDS elemental analysis (the proportion of them), in Fig. 6, and we observed them from microstructural point of view as light grey compact particles, Fig. 7(a) and (d), respectively and Fig. 8. We can notice the graphite (C) as dark grey compact particles in all ESEM microstructure and confirmed by X-ray mapping (Ck) and EDS elemental analysis. Due to the combined influences of powder mixtures components, it is not possible to determine precisely the influence of each component on the basic structure. The most pregnant structural modifications are determined especially by the copper additives. The equilibrium diagram of Fe–Cu shows that copper is liquid at 1083 ◦ C and can easily infiltrate between the Fe particles. There is a partial diffusion in iron and, thus, the net is reinforced (liquid copper ensures a better connection between alloy particles and resulted that by action of capillary force). The above described phenomena can take place at lower temperatures than the melting points of copper, due to the iron-solubility of 8.5% at 1094 ◦ C and of 4.5% at 833 ◦ C.
4.1. Powders and frictional material characteristics
4.2. Influence of braking parameters on friction coefficients
In Table 2 the physical characteristics of raw powders are presented and in Fig. 4 the morphology of elemental powders is presented. We observed in Fig. 4 that iron powders have spherical and egg-shape morphology, the dendrites particle shape is specific for electrolytic copper powders, the mullite has irregular form and the
The braking parameters used were number of braking stops (25–500) and speed of 300–1200 rpm. The friction coefficient variation depending on temperature, the speed and number of brakes for test samples on the inertial stand is presented in Fig. 9. The values obtained by simulating the conditions of real friction, on the inertial test-stand, were brought close
Fig. 3. Sintered friction materials (second lot), fixed on a metallic pad, for test samples on inertial stand.
In this design, n = 2 (i.e. applied load and relative speed in braking point) and P = 2 (i.e. upper and lower level of each variable). Thus, the minimum number of trial experiments to be conducted for each material is four. If wear rate is represented by Y1 and Y2 (Y1 measured by height loss of material and Y2 measured by weight loss) and friction coefficient is represented by Y3 , the linear regression equation for these experiments could be written as: Yi = bi0 + bi1x1 + bi2x2 + bi3x1x2
(5)
where bi0 is the response variable of wear rate and friction coefficient, respectively, at the base level (i.e. at the applied load 7.45 × 10−1 MPa and relative speed in braking points 14.7 m/s); bi1 and bi2 are coefficients associated with each variable x1 (applied load) and x2 (relative speed in braking point) and bi3 is interaction coefficient between x1 and x2 , within the selected levels of each of the variables. The parameters of Eq. (5) have been estimated by the method of least squares using a Excel programme package. The coefficients were determined statistically using the Student criterion and the accordance degree of the regression equations was verified with the Fischer criterion. The mathematical model of wear behaviour once established, the problem of the models optimization is to determination of the mathematical function extreme [30].
Table 2 The physical characteristics of raw powders. Powder type
Fe Cu Graphite MoS2 Mullite (3Al2 O3 .2SiO2 ) Spinel (MgAl2 O4 )
Particle size distribution, % >100 m
100–63 m
63–50 m
50–40 m
<40 m
23.4 – – – – –
28.1 54.4 69.8 44 – –
28.3 24.7 20.4 29 23 36
11.9 12.2 5.1 12.5 42 32.1
8.3 8.7 4.7 14.5 35 31.9
Apparent density, g/cm3
Flow rate, s/50 g
2.55 1.34 0.49 0.43 0.72 0.81
39 41.2 Not flowing Not flowing Not flowing Not flowing
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Fig. 4. Morphology of (a) Fe powder (1000×), (b) Cu powder (1000×), (c) graphite powder (3000×), (d) MoS2 powder (2000×) and (e) Mullite powder (1000×).
to the maximal allowed amount of 0.5–0.6 at the cold testing, under speed increasing conditions (Fig. 9(b)). 4.3. Linear regression equations The upper level and lower level of each variable along with their code values used in these investigations are presented in Table 4. Experiment planning matrix, the values of individual variables with their wear rate and friction coefficient response in each trial tested for SFM1, SFM2 and RM samples (results of random experiments along with theoretical value) are presented in Tables 5–7 (where y1 , y2 and y3 are the experimental values and y1calc , y2calc and y3calc are the calculated ones). The calculation of dispersal reproducibility for y1 , y2 – wear rate and y3 – friction coefficient was made in cases in which the values of Z1 xZ2 have the base level.
The four parallel experiences are presented in Table 8, where u is the number of experience; n is the number of parallel experiments is the arithmetic average of obtained results for k = 1, 2 or (n = 4); yuk 3, the type of variable response; v2 is the degrees of freedom at n − 1. The calculation of reproducibility dispersion is based on the Eq. (6):
n S02
=
u=1
yu
n−1
(6)
Result for variable response Y1 is S02 = 0.00071875, for variable response Y2 is S02 = 0.000583 and for variable response Y3 is S02 = 0.000475, for sample SFM1. The coefficients were determined statistically using the Student criterion and the accordance degree of the regression equations was verified with the Fischer criterion.
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77
Fig. 5. The structure of iron–copper based composites (SFM1): (a) microstructure ESEM, (b) EDS elemental analysis, and (c) compo image of samples, 500× magnification.
Thus, the coefficients of each variable from Eq. (5) were calculated with Eqs. (7) and (8).
n
x · yu u=1 iu , (xiu )2
bi =
n bij =
x u=1 iu
for linear effects,
· xju · yu
(xiu · xju )2
,
for interaction effects,
(7)
(8)
By introducing the calculated coefficients (7) and (8) in Eq. (5) the following regression equations result: Y1,SFM1 = 1.37 + 0.45x1 + 0.07x2 + 0.03x1 x2
(9)
Y2,SFM1 = 0.535 + 0.101x1 + 0.037x2 − 0.0325x1 x2
(10)
Y3,SFM1 = 0.454 − 0.083x1 − 0.0138x1 x2
(11)
where Y1,SFM1 , Y2,SFM1 represent the response variables (wear rate of SFM1 samples measured by height loss and weight loss, respectively; Y3,SFM1 represents the friction coefficients of SFM1). The dispersion in determination of coefficients is calculated with relations: 2 Sbi =
S02
x j=1 ij
and Sbi =
2 Sbi
(12)
Table 4 Levels of each factor and their coded values. SI. no.
1 2 3
Variables Factor levels
Z1 – Applied load (×10−1 MPa)
Z2 – Relative speed in braking point (m/s)
Coded values Upper level (+1) Base level, Z10 (0) Lower level (−1) Variation interval, Z1
x1 11.5 7.5 3.5 4
x2 16.9 14.7 12.5 2.2
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Fig. 6. The X-ray mapping of the elements (a) and EDS elemental analysis of the frictional sintered iron–copper based material (SFM1) (b).
Result for variable response Y1 , Sbi = 0.013404757, for variable response Y2 , Sbi = 0.012076, and variable response Y3 , Sbi = 0.010897. The Student criterion t0.05;4 = 2.776 and bi – confidence interval is bi = 2.776·Sbi . After calculation, result for variable response Y1,SFM1 , bi = 0.0372, for variable response Y2,SFM1 , bi = 0.0335 and for variable response Y3,SFM1 , bi = 0.0302. But according to [30], if the relation b0 , b1 bi is not accomplished, result for the calculated value of the coefficient is not statistically different
from zero; therefore, the term comprising interaction, or another factor from equation, will be neglected. In this case, the calculated (predicted) linear model of the processes will be: for the wear, Eqs. (13) and (14) and friction coefficients Eq. (15), respectively, of the friction material SFM1 can be expressed as follows: Y1,SFM1 = 1.37 + 0.45x1 + 0.07x2
(13)
Y2,SFM1 = 0.535 + 0.101x1 + 0.037x2
(14)
Y3,SFM1 = 0.454 − 0.083x1
(15)
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79
Fig. 7. The microstructure ESEM (SFM1), 500× maginification (a) and EDS elemental analysis taken from the compact white phase (b), from small white particles (c) and light grey compact particles (d).
In this equation x1 and x2 are coded values of the factors, which are calculated from natural values, with the formula: xi =
zi − zio zi
(16)
Substituting in Eq. (16), the values from Table 4, we will z −7.5 z −14.7 obtain(17)x1 = 1 4 and x1 = 2 2.2
Verification of the model concordance: The computation of dis2 persion produced by the linear equations of regression Sconc is presented in Table 9, where
yu = yexp − y
Fig. 8. The microstructure ESEM (SFM1), 2000× maginification (a) and EDS elemental analysis taken from the light grey compact particles (b).
(18)
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Fig. 9. The variation friction coefficient depending on temperature, speed and number of brakes for test samples on the inertial stand in: grinding (a), cold testing (b), hot testing (c), recovery testing (d), and endurance conditions (e).
and n 2 Sconc =
(yu )
2
u=1
1
(19)
The calculated Fischer criterion is: Fc =
2 Sconc
S02
(20)
Introducing the values of yu from Table 9 in Eq. (19) and then 2 introducing the resulted Sconc and values of reproducibility disper2 sion S0 in Eq. (20) result that the calculated Fisher is Fc1 = 5.01 (for variable response Y1 ); Fc2 = 7.307 (for variable response Y2 ) and Fc3 = 5.05 (for variable response Y3 ). From tabulated values of optimization book annexes in [29] resulted in F˛ , 1 , 2 = F0.05,1.3 = 10.13. Because Fci < F0.05,1.3 , i = 1.2 or 3 ⇒. The linear models for all three variables of sintered friction material SFM1 are in concordance with analyzed process.
Table 5 The values of individual variables with their wear rate and friction coefficient response in each trial tested for SFM1 samples. Trial no.
1 2 3 4
x0
Applied load (×10−1 MPa) x1
Relative speed (m/s) x2
(+1) (+1) (+1) (+1)
11.5 (+1) 3.5 (−1) 11.5 (+1) 3.5 (−1)
16.9 (+1) 16.9 (+1) 12.5 (−1) 12.5 (−1)
x1 x2
Wear rate (mm/h) y1
Calculated wear rate (mm/h) y1calc
Wear rate (g/cm2 h) y2
Calculated wear rate (g/cm2 h) y2calc
Friction coefficient y3
Calculated friction coefficient y3calc
(+1) (−1) (−1) (+1)
1.92 0.96 1.72 0.88
1.89 0.99 1.75 0.85
0.64 0.503 0.631 0.364
0.673 0.471 0.599 0.397
0.337 0.53 0.405 0.543
0.3709 0.537 0.371 0.537 I.N. Popescu et al. / Applied Surface Science 285P (2013) 72–85
Table 6 The values of individual variables with their wear and friction coefficient response in each trial tested for SFM2 samples. Trial no.
1 2 3 4
x0
Applied load (×10−1 MPa) x1
Relative speed (m/s) x2
(+1) (+1) (+1) (+1)
11.5 (+1) 3.5 (−1) 11.5 (+1) 3.5 (−1)
16.9 (+1) 16.9 (+1) 12.5 (−1) 12.5 (−1)
x1 x2
Wear rate (mm/h) y1
Calculated wear rate (mm/h) y1calc
Wear rate (g/cm2 h) y2
Calculated wear rate (g/cm2 h) y2calc
Friction coefficient y3
Calculated friction coefficient y3calc
(+1) (−1) (−1) (+1)
1.4 1.1 1.25 0.86
1.424 1.078 1.228 0.882
0.52 0.29 0.368 0.26
0.491 0.321 0.399 0.229
0.295 0.395 0.32 0.46
0.295 0.395 0.32 0.46
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0.353 0.529 0.287 0.463 0.39 0.49 0.25 0.5 0.723 0.317 0.759 0.353 0.65 0.39 0.83 0.28 1.559 1.333 1.433 1.207 1.56 1.33 1.43 1.21 (+1) (−1) (−1) (+1) 16.9 (+1) 16.9 (+1) 12.5 (−1) 12.5 (−1) 11.5 (+1) 3.5 (−1) 11.5 (+1) 3.5 (−1) (+1) (+1) (+1) (+1)
x1 x2 x0
In the same way the 4 parallel experiences for variable response for SFM2 and RM samples were calculated and the following linear regression equations (21)–(26) were obtained which are also in concordance with analyzed process:
1 2 3 4
Calculated wear rate (mm/h) y1calc Wear rate (mm/h) y1 Relative speed (m/s) x2 Applied load (×10−1 MPa) x1 Trial no.
Table 7 The values of individual variables with their wear and friction coefficient response in each trial tested for RM samples.
Wear rate (g/cm2 h) y2
Calculated wear rate (g/cm2 h) y2calc
Friction coefficient y3
Calculated friction coefficient y3calc
82
Y1,SFM2 = 1.153 + 0.173x1 + 0.098x2
(21)
Y2,SFM2 = 0.359 + 0.085x1 + 0.046x2
(22)
Y3,SFM2 = 0.368 − 0.06x1
(23)
Respectively: Y1,MR = 1.383 + 0.113x1 + 0.063x2
(24)
Y2,MR = 0.538 + 0.203x1 − 0.073x1 x2
(25)
Y3,MR = 0.408 − 0.088x1
(26)
From Eq. (21) deduced that the height loss increase with increasing of applied load, relative speed in braking point and no depends with interaction of those variables. Eq. (22) showed that the weight loss of SFM2 material depends direct proportional with applied load and no depend of relative speed (witch means that regardless of automotive speed, braking to accomplish in safe condition depending on pressing force. Eq. (23) showed that the friction coefficient of SFM2 material increased with increased all the variables inclusive with interaction of them. We noticed that wear rate measured by height loss for reference material (Y1,MR ), weight loss of material increased with increasing of applied load and relative speed in accordance with Eq. (24) and wear rate measured by weight loss (Y2,MR ) increased with increasing of applied load and with decreasing of interaction of applied load and relative speed, in accordance with Eq. (25). Eq. (26) shows that the friction coefficient of reference material decreases with increasing of applied load. The calculated (theoretical) values of wear rate and friction coefficient were obtained replacing in Eqs. (13)–(15) and (21)–(26) the coded values corresponding of each trial for all types of material tested. Through the statistical experimental design presented before, we can describe the evolution of tribological characteristics of sintered frictions samples (SFM1, SFM2 and RM) at different applied load, different relative speed and constant temperature. Thus, for comparison, the statistical results of response variables, represented by wear rates and friction coefficients, and the experimental data of them, are presented in Figs. 10 and 11. For better understanding of tribological behaviour of the three types of samples depending on all factors: applied load (3.5 × 10−1 MPa and 11.5 × 10−1 MPa) and relative speed (16.9 m/s and 12.5 m/s), we presented all the results of experimental and calculated data in Fig. 11. From these figures we observed that the difference between experimental values and the calculated one are the majority ±3–6%. Comparing the wear rates of SFM1 and SFM2 samples with reference material RM we noticed that values for all materials are not great; in addition at maximum values of applied load and relative speed, wear rate measured by loosing weight per hour is 0.64 g/cm2 h for SFM1, 0.32 g/cm2 h for SFM2, 0.65 g/cm2 h for MR, and wear rate measured by loosing height per hour is 1.92 mm/h for SFM1, 1.4 mm/h for SFM2 and 1.56 mm/h for RM. We have seen that the lowest values of wear rate have friction materials SFM2, which have a lowest value in comparison with the reference material. The values of friction coefficients at maximum solicitations are:
I.N. Popescu et al. / Applied Surface Science 285P (2013) 72–85 Table 8 The 4 parallel experiences for variable response for SFM1 sample.
No.
y ’ u1
yu1 = y u1 − yu1
(yu1 )
1 2 3 4
1.52 1.6 1.64 1.43
0.0275 0.0525 0.0925 0.1175
0.00075625 0.00275625 0.00855625 0.01380625
2
y ’ u2
yu2 = y u2 − yu2
(yu2 )
0.51 0.48 0.59 0.5
0.01 0.04 0.07 0.02
0.0001 0.0016 0.0049 0.0004
2
83
y’3
yu3 = y u3 − yu3
(yu3 )
0.47 0.42 0.47 0.38
0.035 0.015 0.035 0.055
0.001225 0.000225 0.001225 0.003025
2
2 3 3 3 3
Table 9 2 produced by the linear equation of regression, for SFM1 sample. The computation of dispersion Sconc No.
y1exp
y¯ 1
y1u
(y1u )2
y2exp
y¯ 2
y2u
(y2u )2
y3exp
y¯ 3
y3u
(y3u )2
1 = N − k
1 2 3 4
1.92 0.96 1.72 0.88
1.89 0.99 1.75 0.85
0.03 0.03 0.03 0.03
0.0009 0.0009 0.0009 0.0009
0.64 0.503 0.631 0.364
0.673 0.471 0.599 0.397
0.033 0.032 0.032 0.033
0.001089 0.001024 0.001024 0.001089
0.337 0.53 0.405 0.543
0.371 0.537 0.371 0.537
0.034 0.007 0.034 0.006
0.001156 0.000049 0.001156 0.000036
1 1 1 1
0.337 for SFM1, 0.36 for SFM2, 0.39 for MR respectively. We noticed that SFM1 has a constant and higher value of friction coefficient; this value is closer to a coefficient of friction value of MR material. From Figs. 10 and 11, we can see that the values of friction coefficients have a low decrease with increasing of strains (load and speed), decreasing from 0.543 to 0.337 for SFM1, from 0.46 to 0.295for SFM2, from 0.5 to 0.39 for RM (maximum value was obtained for applied load of 3.5 × 10−1 MPa and 16.9 m/s speed);
all values of frictional coefficients are in admitted domain (0.5–0.3) for automobile friction materials. From statically point of view, after we analyzed Eqs. (13)–(15) for SFM1 sintered friction material, Eqs. (21)–(23) for SFM2 sintered friction material and Eqs. (24)–(26) for reference material RM respectively were found them “in accordance” on the basis of the Fischer criterion which means that the model is good and supports the reproducibility of the trial experiments.
Fig. 10. The experimental and statistical calculated values of response variables: (a) wear rates of samples as a function of applied load, at constant speed (16.9 m/s), (b) friction coefficients of samples as a function of applied load, at constant speed (16.9 m/s), (c) wear rates of samples as a function of applied load, at constant speed (12.5 m/s); (d) friction coefficients of samples as a function of applied load, at constant speed (12.5 m/s).
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Fig. 11. The experimental and calculated values of wear rate and friction coefficient respectively of all three types of materials as a function of applied load (3.5, 11.5 × 10−1 MPa, respectively), for relative speed (a) 16.9 m/s and (b) 12.5 m/s.
5. Conclusions We obtained a sintered friction iron–copper based material, with new chemical compositions and proportion of components (SFM1 and SFM2 samples) and non-polluting constituents (ecologically friendly). The chosen composition and proportion of these components ensured good physical and mechanical characteristics, confirmed also by the morphological and microcompositional investigation (uniform distribution of constituents and good adhesion between them, due to changes occurring during liquid phase sintering in the Fe–Cu system). We established after different conditions of tribological tests, including the tests on inertial stand (inertia dynamometer) which simulate the real conditions of dry friction, that the obtained frictional materials for braking pad have: (i) high and stable friction coefficients (0.3–0.5), according to the requirements imposed by brake pads materials for automotive application (Fig. 9); (ii) high durability of SFM2 materials in comparison with RM and SFM1 samples; the values of wear rate, in the same experimental conditions for SFM2 is maximum 25% lower than RM (Figs. 10 and 11). (iii) increase of applied load or increase of the relative speed in braking points induces moderate increase of wear rate and slow decrease of the frictional coefficients of the composites (shown through experimental data and obtained linear regression equations in Figs. 9–11). Acknowledgements The authors are grateful to the Metallurgical Research Institute Bucharest and BIOMAT Research Centre, Bucharest for providing equipments and also thank Full Professor Ph.D. Dionezie Bojin from Electron Microscopy Laboratory, Science and Engineering Faculty, POLITEHNICA University of Bucharest for advice and support. References [1] H. Jang, K. Ko, S.J. Kim, R.H. Basch, J.W. Fash, The effect of metal fibers on the friction performance of automotive brake friction materials, Wear 256 (3–4) (2003) 406–414.
[2] H. Zaidi, A. Senouci, Thermal tribological behaviour of composite carbon metal/steel brake, Applied Surface Science 144–145 (1999) 265–271. [3] N. Popescu, E. Geza, Contribution to increase of the performant composite materials having friction properties, in: Proceedings of the 2nd National Powder Metallurgy Conference, Metallurgical and Materials Engineering Dept., Ankara, Turkey, 15–17 July, 1999, pp. 225–233. [4] P.J. Blau, H.M. Meyer III, Characteristics of wear particles produced during friction tests of conventional and unconventional disc brake materials, Wear 255 (2003) 1261–1269. [5] M. Eriksson, F. Bergman, S. Jacobson, Surface characterization of brake pads after running under silent and squealing conditions, Wear 232 (1999) 163–167. [6] Y.M. Wang, B.L. Jiang, L.X. Guo, T.Q. Lei, Tribological behavior of microarc oxidation coatings formed on titanium alloys against steel in dry and solid lubrication sliding, Applied Surface Science 252 (2006) 2989–2998. [7] A.E. Anderson, Friction and wear of automotive brakes, in: ASM handbook, vol. 7. Powder Metallurgy, ASM International, USA, 1984, pp. 569–577. [8] D. Chan, G.W. Stachowiak, Review of automotive brake friction materials, Proceedings of the Institution of Mechanical Engineers, Part D: Journal of Automobile Engineering 218 (2004) 953–966. [9] M.L. Marucci, HanejkoF F.G., Effect of Copper Alloy Addition Method on the Dimensional Response of Sintered Fe–Cu–C Steels, 2010, pp. 1–11 http://www.mpif.org [10] T. Murphy, M. Baran, An investigation into the effect of copper and graphite additions to sinter hardening steels, in: Advances in Powder Metallurgy and Particulate Materials, part 10, Metal Powder Industries Federation, Princeton, NJ, 2004, pp. 266–274. [11] M. Morakotjinda, R. Krataitong, P. Wila, P. Siriphol, A. Daraphan, O. Coo, Sintered Fe-Cu-C materials, Chiang Mai Journal of Science 35 (2) (2008) 258–265. [12] V. Prasad, P.K. Rohatagi, Tribological properties of al alloy particle composite, Journal of Metals 11 (1987) 22. [13] X. Xiong, H-C. Sheng, C. Jie, Y. Ping-ping, Effects of sintering pressure and temperature on microstructure and tribological characteristic of cu-based aircraft brake material, Transactions of Nonferrous Metals Society of China 17 (2007) 669–675. [14] E. Hounkponou, H. Nery, D. Paulmier, A. Bouchoucha, H. Zaidi, Tribological behaviour of graphite/graphite and graphite/copper couples in sliding electrical contact: influence on the contact electric field on the surface passivation, Applied Surface Science 70–71 (1) (1993) 176–217. [15] Popescu IN, Ernszt G. Researches Projects Orizont 2000 No. C 55B, Theme A33/3. Composites materials with iron based matrix used for frictional materials (in Romanian). [16] I.N. Popescu, S. Zamfir, V.F. Anghelina, C.O. Rusanescu, Processing by P/M route and characterization of new ecological aluminum matrix composites (AMC), International Journal of Mechanics 3 (4) (2010) 71–80. [17] I.N. Popescu, S. Zamfir, D. Bojin, M. Brânzei, D. Gheorghe, L.A. Sorcoi, Physical and quantitative microstructural analysis of sintered Al–Cu/SiCp composites, Materials Science Forum 672 (2011) 251–254. [18] S. Chen, C. Kang, J. Wang, C. Liu, K. Sun, Synthesis of anodizing composite films containing superfine Al2 O3 and PTFE particles on Al alloys, Applied Surface Science 256 (2010) 6518–6525. [19] V. Tsakiris, W. Kappel, E. Enescu, G. Alecu, F. Albu, F. Grigore, V. Marinescu, M. Lungu, Characterization of Al matrix composites reinforced with
I.N. Popescu et al. / Applied Surface Science 285P (2013) 72–85
[20]
[21]
[22]
[23]
[24]
alumina nanoparticles obtained by pm method, Journal of Optoelectronics and Advanced Materials 13 (9) (2011) 1172. M. Lucaci, M. Valeanu, R.L. Orban, V. Tsakiris, D.C. Cristea, L. Leonat, Shape memory NITI alloys obtained by powder metallurgy route, Materials Science Forum 672 (2011) 99–104. V. Andrei, E. Andrei, G.H. Vlaicu, C. Stihi, G. Dima, C. Oros, S. Dinu, Chemical binding and structure of carbonic thin films with advanced properties studied by electron spectroscopy, Journal of Optoelectronics and Advanced Materials 9 (7) (2007) 2291–2295. I.N. Popescu, D. Bojin, I. Carceanu, G. Novac, F.V. Anghelina, Morphological and structural aspects using electronic microscopy and image analysis of iron powders obtained by water atomization process, Journal of Science and Arts 1 (12) (2010) 125–134. S. Domsa, M. Firanescu, Influence of sintering conditions on properties of sintered friction materials based on iron powder, Metallurgy and New Materials Researches 18 (1997) 655. R.I. Zamfir, The influence of the laser surface melting parameters (LSM) on the structure and microhardness of the 316L austenitic stainless steel, UPB Scientific Bulletin 71 (2009) 137.
85
[25] J. Hernáez, E. Otero, A. Pardo, C. Merino, M. Laguna, Contribution to the microstructural study of Fe-8% Cu sintered compacts, Revista de Metalurgia (Madrid) 22 (2) (1986) 95–101. [26] D. Bojin, D. Bunea, Fl. Minculescu, M. Minculescu, Scanning Electron Microscopy and Applications, AGIR, Bucharest, 2005 (in Romanian). [27] Y. Sahin, Wear behavior of aluminium alloy and its composites reinforced by SiC particles using statistical analysis, Materials & Design 24 (2003) 95–103. [28] I.N. Popescu, V. Bratu, M.C. Enescu, Experimental researches and statistical analysis of the corrosion behavior of rolled and heat treated 2XXX Al alloys, in: AEE’10 Proceedings of the 9th WSEAS, International Conference on Applications of Electrical Engineering, Recent Advances in Electrical Engineering (2010) 225–232. [29] C. Ghit¸a˘ , Mathematical Metallurgy, Romanian Academy, Bucharest, 1995 (in Romanian). [30] D. Taloi, E. Florian, C. Bratu, E. Berceanu (Eds.), The Optimisation of Metallurgical Processes, Didactica & Pedagogic˘a, Bucharest, 1983 (in Romanian).