S Gustafsson et al. /Joumal of the European Ceramic Sociery 29(2009)539-550 8 gm( 2um Fig. 2. Thermally etched surfaces of the polycrystalline mullite in the(a)as-sintered condition, and after creep testing under a stress of ( b)48.6 MPa at 1400C,(c) 130MPat1400°C,and(d)149 MPa at1300°C showed that the stress exponent increased with increasing stress; 3. Experimental procedures om around n=1.5 at stresses under 25 MPa to around n=4 at stresses above 25 MPa. This implies that the total strain was not 3.1. Materials caused by one single creep mechanism The creep rate intervals for diffusion-controlled creep of 3.1.1. Polycrystalline mullite polycrystalline mullite with a reduced average grain size The polycrystalline mullite material was produced by mxIn =0.7 um, corresponding to the average matrix grain size commercially available 3: 2 mullite powder(KM-10l, Kyoritsu, of the nanocomposite)were calculated as described in Sec- Japan) and an ammonium polyacrylate dispersant (Dispex tion 2. 1. 8,9 This was done in order to better assess the Allied Colloids, England) in water. The slurry was ball milled effect of the SiC particles, and these creep rate intervals for 24 h using zirconia ball milling beads. Green bodies were re also shown in Fig. 1b. As illustrated in Fig. Ib, the produced by slip casting and pressureless sintered in air at experimental creep rates of the nanocomposite tested at low 1650C for 3 h. The material was 97% dense as measured by stresses(<30 MPa)at 1400C were lower than the pre- Archimedean densitometry dicted diffusion creep rates of polycrystalline mullite with this grain size. At higher stresses, however, the creep rate of the 3. 1.2. Mullite/Sic nanocomposite predicted by the diffusion- The a-SiC starting powder (UF-45, H.C. Starck, Germany) controlled creep model. The two data points from creep tests had a specific surface area of around 45 m-/g. The larger particles at 1300C were within the predicted diffusion creep rate inter- and agglomerates that were difficult to break down by milling powder were removed by sedimentation. This resulted in nanocomposite was determined not only by the reduced mullite particle size(dso)of 0.22 will starting powder that had a mean drive self-diffusion in the low diffusivity SiC particles, so that aqueous suspension containing 95 vol. of the mullite pm grain size. It has been suggested that the extra work required to The nanocomposite material was then produced from they can move with the grain boundaries during creep, will lead der, 5 vol. of the milled and fractionated a-SiC powder, to a reduced creep rate as compared to polycrystalline mullite and 0.3 wt% of an ammonium polyacrylate dispersant(Dura of the same grain size max 3021, Rohm and Haas, Sweden). The suspension wasS. Gustafsson et al. / Journal of the European Ceramic Society 29 (2009) 539–550 541 Fig. 2. Thermally etched surfaces of the polycrystalline mullite in the (a) as-sintered condition, and after creep testing under a stress of (b) 48.6 MPa at 1400 ◦C, (c) 13.0 MPa at 1400 ◦C, and (d) 14.9 MPa at 1300 ◦C. showed that the stress exponent increased with increasing stress; from around n = 1.5 at stresses under 25 MPa to around n = 4 at stresses above 25 MPa. This implies that the total strain was not caused by one single creep mechanism. The creep rate intervals for diffusion-controlled creep of polycrystalline mullite with a reduced average grain size (d = 0.7m, corresponding to the average matrix grain size of the nanocomposite) were calculated as described in Sec￾tion 2.1. 18,19 This was done in order to better assess the effect of the SiC particles, and these creep rate intervals are also shown in Fig. 1b. As illustrated in Fig. 1b, the experimental creep rates of the nanocomposite tested at low stresses (<30 MPa) at 1400 ◦C were lower than the pre￾dicted diffusion creep rates of polycrystalline mullite with this grain size. At higher stresses, however, the creep rate of the nanocomposite was in the range predicted by the diffusion￾controlled creep model. The two data points from creep tests at 1300 ◦C were within the predicted diffusion creep rate inter￾val. The data presented in Fig. 1b indicate that the creep rate of the nanocomposite was determined not only by the reduced mullite grain size. It has been suggested that the extra work required to drive self-diffusion in the low diffusivity SiC particles, so that they can move with the grain boundaries during creep, will lead to a reduced creep rate as compared to polycrystalline mullite of the same grain size.19 3. Experimental procedures 3.1. Materials 3.1.1. Polycrystalline mullite The polycrystalline mullite material was produced by mixing commercially available 3:2 mullite powder (KM-101, Kyoritsu, Japan) and an ammonium polyacrylate dispersant (Dispex, Allied Colloids, England) in water. The slurry was ball milled for 24 h using zirconia ball milling beads. Green bodies were produced by slip casting and pressureless sintered in air at 1650 ◦C for 3 h. The material was 97% dense as measured by Archimedean densitometry. 3.1.2. Mullite/SiC nanocomposite The -SiC starting powder (UF-45, H.C. Starck, Germany) had a specific surface area of around 45 m2/g. The larger particles and agglomerates that were difficult to break down by milling the powder were removed by sedimentation. This resulted in a milled and fractionated SiC starting powder that had a mean particle size (d50) of 0.22m. The nanocomposite material was then produced from an aqueous suspension containing 95 vol.% of the mullite pow￾der, 5 vol.% of the milled and fractionated -SiC powder, and 0.3 wt% of an ammonium polyacrylate dispersant (Dura￾max 3021, Rohm and Haas, Sweden). The suspension was