We report for the first time use of 3D-printed metal extrusion dies for the extrusion of an optical material into an extrudate at elevated temperatures. Using lead-silicate glass as the material to be extruded, the 3D-printed dies demonstrated the same glass flow behavior as conventionally machined metal dies. Evaluation of the extrusion force at set temperature and extrusion speed revealed that the metal-type of the dies used did not affect the glass flow behavior. Using 3D-printed dies as delivered, the high surface roughness of the 3D-printed dies resulted in high preform surface roughness. However, this effect was overcome by finishing the easily accessible internal die surfaces over 1-2mm length upstream from the die exit. The opportunity of using 3D-printed dies offers unprecedented flexibility in the die design for unlimited tailoring of fluid flow within the die, which paves the way towards extruded items of arbitrary shape.
© 2014 Optical Society of America
Extrusion of optical glasses and polymer billets has been demonstrated to be a technique suitable for the fabrication of fiber preforms with a wide range of structures [1–39]. In the billet extrusion technique, bulk billets are heated up to a temperature where the optical material is sufficiently soft to be forced through a die structure by a ram to form an extrudate, which has a transverse profile with millimeter-scale features that is determined by the die exit geometry . Billet extrusion requires viscosities in the range of 106-1010 Pa.s. At lower viscosities, the soft material emerging from the die exit deforms severely due to gravity and surface tension. At higher viscosities, the material is not sufficiently soft to be forced through the die without material damage such as cracks and/or without die damage such as surface abrasion, leading to preform contamination, and even die breakage [2, 3]. To fabricate low-loss optical fibers, the starting preforms need to have good optical quality. Extrusion of single-material billets into preforms with high optical quality is usually performed at relatively low ram speeds of 0.02-0.2mm/min to avoid material damage and die surface abrasion [2, 3]. Both effects severely deteriorate the optical quality of preforms, which can lead to undesirable increased loss for the fibers made from such preforms . For multimaterial co-extrusion of chalcogenide glass and polymer, higher ram speeds of 0.3-0.7 mm/min are used to protect the chalcogenide glasses thermally and mechanically .
The extrusion technique offers several advantages over other preform fabrication techniques such as stacking, drilling and casting. Extrusion is a computer-controlled, automated process, which enables a high degree of reproducibility in contrast to the manual fabrication of stacked and cast preforms. Compared to preforms made by drilling holes into rods, extrusion offers the benefit of leading to smooth surfaces during the preform forming process due to the fire-polishing effect at elevated temperatures, i.e. at low glass viscosities. The glass preform surface is still hot and thus deformable when the preform emerges from the die exit and moves slowly out of the hot zone. The deformability of the surface at low viscosity enables rounding of sharp edges and filling of scratches on the surface . In contrast, high surface quality of drilled preforms can only be achieved after the preform forming process in a post-polishing step, which adds time to the preform fabrication and, for small and long holes, can be challenging to be undertaken.
The extrusion can be applied to any material that does not degrade in the viscosity range that can be used for billet extrusion. To date, billet extrusion has been demonstrated for a wide range of optical materials; polymer (polymethymethacrylate [5–7], cyclic olefin copolymer (TOPAS) ), various lead-silicate glasses [2, 8–20], tellurite glasses [21–25], bismuth glasses [26, 27], germanate glass , phosphate glass , chalcogenide glasses [4, 30–34], various fluoride glass types [3, 35–38] and fluoride-phosphate glass . For all these materials, the softening temperature is lower than 650 °C. The extrusion of glasses with higher softening temperatures is not limited by the glass material itself but by the availability of die material and heating apparatus. To date, stainless steel dies have been most widely used as extrusion die material. It exhibits high mechanical strength at elevated temperatures up to ~650 °C, can be readily machined into complex structures, is cost-effective, and does not show surface abrasion against flowing glass at the elevated temperatures using the glass viscosity and ram speed ranges mentioned above.
Another advantage of extrusion is that all features in a preform are made simultaneously, which is in contrast to stacking and drilling where the holes are created sequentially. A further benefit of extrusion is the potential to fabricate any structure provided a die structure can be designed that enables feasible glass flow in the practical range of glass viscosity and ram speeds described above.
To date, the potential of extrusion to fabricate any arbitrary shape has been mostly used to create non-circular holes and outer shapes [5–11, 14–19, 21, 22, 24, 26–28]. For a few preforms, non-rotational symmetric hole patterns have also been demonstrated [6, 7, 17]. The extrusion of preforms with larger degree of non-circular and non-rotational symmetric features and patterns is limited by constraints associated with the machining of the extrusion dies. Current dies for extruding complex preforms exhibit mostly circular feed holes and blockages as such features can be easily machined using traditional machining techniques. While other feature shapes can be made in principle, the machining of such dies is cost-prohibitive. The advent of 3D printing technology overcomes these die machining restrictions. 3D printing allows fabrication of almost any arbitrary shapes in three dimensions with high precision [40, 41], including shapes that cannot be made using conventional subtractive machining techniques such as conformal cooling channels . Recently, 3D printing has been developed for a range of metals, which opens up the possibility of producing 3D-printed extrusion dies that can withstand extrusion forces and elevated temperatures. The geometrical freedom offered by 3D printing has been exploited for the fabrication of an aluminium extrusion die exhibiting conformal cooling channels while the structure of the die was not weakened .
In this paper, we report the use of 3D-printed extrusion dies for extrusion of lead-silicate glass. To the best of our knowledge, this is the first demonstration that 3D-printed dies are a viable alternative to dies made using conventional, subtractive machining methods. We evaluated the performance of the 3D-printed dies by comparing glass flow behavior through the 3D and conventionally machined dies. It is known that 3D-printed metal items can have a relatively high surface roughness in the order of tens of micrometers . Thus, the question arises whether the high surface roughness of 3D-printed dies (compared to conventionally machined dies) leads to preforms with relatively high surface roughness, which ultimately can increase loss of the optical fibers drawn from such preforms. Therefore, we also investigated the impact of die surface roughness on the surface quality of the preforms made using both 3D-printed and conventionally machined dies.
2. Extrusion experiments
We selected the lead-silicate glass F2 from Schott Glass Co for the extrusion experiments as this glass has been widely used to extrude various structures [14–19], and the glass flow behavior through machined dies is well characterized [2, 20]. The extrusion temperature of 560-600 °C is relatively high, allowing 3D-printed dies to be tested at relatively harsh extrusion conditions. To enable comparison with machined dies, we selected a relatively simple die structure for the extrusion of a tube with 10mm outer diameter and 2mm inner diameter. The design of the dies is based on our truly scalable die design concept that allows the definition of a large number of transverse features . These dies exhibit two main parts: (i) a traverse with an array of feed holes going through the traverse and with an array of blockages protruding from the traverse, and (ii) a welding chamber defining the outer shape of the preform, with the blockages being situated within the welding chamber (Fig. 1). In the extrusion process, first the glass flows through the feed holes into the welding chamber. The individual glass strands emerging from the feed holes fuse together into a single body in the welding chamber, while the blockages obstruct the flow and thus form the holes in the preform. The die used here to make preforms with a tube structure contain one blockage to form the tube hole. Note that this scalable die requires undercuts around thin blockages to enable sufficient mechanical strength of the traverse. Due to these undercuts, conventional machining requires the traverse with feed holes and blockage to be made separately from the die outer defining the welding chamber. In contrast, the 3D-printed dies were made in a single piece, which is of particular importance for dies that rely on correct alignment of inner die features relative to a non-circular die exit. Manufacture of dies composed of a single piece overcomes the limitation of manual alignment or introduction of additional design features to ensure precise alignment for such dies.
We conducted five extrusion trials using 3D-printed and conventionally machined dies (Table 1). We acquired 3D-printed dies made from three different metals: a Co-Cr-Mo superalloy, a tool steel and a Ti alloy [41, 43]. The Co-Cr-superalloy and tool steel die had slightly different die design compared to the Ti alloy die (trials A-C with designs I and II in Table 1). We used two types of machined dies (trials D-E with designs II and III in Table 1) to allow investigation of the impact of die design and extrusion force as detailed below. All five extrusion trials were conducted using the same fixed extrusion process parameters (temperature and ram speed). The die properties and extrusion conditions of the five extrusion trials considered here are summarized in Table 1.
The extrusion trials were conducted as follows. First the extrusion body holding the 3cm diameter billet and the die was heated up to 565 °C, and the temperature was controlled by a thermocouple placed inside the extrusion container. The die temperature was measured via thermocouples placed in contact with the outside of the die near the die exit. When the extrusion assembly including billet and die was at the set temperature, the ram was set to move down at a constant speed of 0.2 mm/min, thus forcing the glass through the die. The electronic feedback control loop of the extrusion machine was set to automatically adjust the ram force to maintain a set ram speed. After the die was filled with glass, and glass starts to emerge from the die exit, the glass flow enters a steady-state regime, where the force did not change significantly with time and ram position. For the five trials considered here, the force in this steady-state regime was in the range of 3-8 kN. The reasons for the extrusion trials yielding different force values are discussed below. Towards the end of extrusion, the force increased gradually up to 20 kN, which was the set force limit. Even at these high extrusion forces, no die breakage was observed.
The force in the steady-state regime was used for the Poiseuille-type glass flow calculations of flow restriction within the dies as described in Section 3. The temperature of the glass within the die was assumed to be equal to the measured die temperature. The viscosity was calculated from the glass temperature using the Arrhenius equation and coefficients for F2 glass . The details of the extrusion set-up and experimental errors are described in . The set and variable (measured) extrusion process parameters are listed in Table 1.
The surface finish of 3D printed items can be very dependent of the type of 3D printer and printer settings, and as such, items can exhibit a higher surface roughness when compared with machined items. Therefore, 3D-printed metal items are usually treated to improve the surface finish. Surface finishing can include abrasive media blasting, machining and hand polishing. The 3D-printed dies used in this work were media blasted by the manufacturers. For the 3D-printed die of trial C, the manufacturer estimated the surface roughness of the die out of the 3D printer and after media blasting to be in the order of 15-20 µm and 10µm, respectively. For the supplied die of trial C, the internal surfaces were finished at the die exit (Fig. 1) using a lathe, whereby a maximum of 0.05 ± 0.01 mm material was removed from the surface. Note that only the easily accessible internal surfaces at the die exit extending over short length of 1-2mm upstream into the die were finished. Note that no polishing was used to treat any surfaces of both 3D-printed and machined dies. The low loss achieved for fibers made from preforms extruded through machined dies without polished surfaces  demonstrates that polishing is not necessary for machined steel dies. All dies were cleaned ultrasonically in detergent prior to use to remove any grease and loose particles.
3. Glass flow analysis
We recently demonstrated that F2 glass extrusion through machined stainless steel dies can be described using the Poiseuille-type flow for isothermal laminar flow of an incompressible viscous liquid with no slip at the glass/die boundary. The Poiseuille-type flow is characterized by the extrusion force being proportional to the glass viscosity, billet diameter and ram speed, with the proportionality constant being the die constant that is determined by the die geometry. For the same die geometry, the Poiseuille flow thus predicts that the die constant calculated using the extrusion processing parameters (billet, diameter, viscosity, ram speed) is constant. To test whether extrusion through 3D-printed dies also obey Poiseuille flow, we calculated the die constants for the five trials considered here (Fig. 2, Table 1).
The 3D-printed dies used for trials A-C exhibited two die designs (I, II) that differed slightly in the welding chamber diameter and the top diameter of the blockage (Fig. 1). The machined die of trial D exhibited the same design as for the 3D-printed die of trial C. The machined die of trial E exhibited a different die design III compared to the designs I and II.
For the same die design (trials A, B with design I and C, D with design II), the same experimentally determined die constants were obtained (Fig. 2, Table 1). The difference in the die constant between the two die designs scales with the cross-sectional area of the welding chamber (Fig. 2). According to the Poiseuille flow, the smaller the cross-sectional area of flow, the larger the die constant. The results for the die constant demonstrate that the 3D-printed dies do not significantly impact the glass flow behavior. In addition, the results indicate that the four die materials used led to similar glass flow behavior.
Considering the two machined dies of trials D and E, the different die design of trial E resulted in comparable force relative to the 3D-printed die trials A and B, whereas 3D-printed die trial C and machined die trial D exhibited similar extrusion force. This data set allowed investigation of the impact of force on preform surface roughness as detailed in the next Section.
4. Preform surface quality
For fluoride and fluoride-phosphate glass, the surface roughness of extruded preforms was shown to affect the loss of the fibers made from these preforms [3, 39]. 3D-printed dies can have a high surface roughness, which could adversely affect the surface quality of extruded preforms. As it can be difficult to improve the surface finish of all internal surfaces of our dies having 18 mm length, we tested whether improving the finish of the easily accessible internal surfaces extending over short length of 1-2mm from the die exit (Fig. 1) was sufficient to achieve preforms with high surface quality. As described in Section 2, the internal die exit surfaces of the 3D-printed die of trial C were finished using a conventional machining tool. Using this approach all of the other internal surfaces remain relatively rough.
To quantitatively determine the impact of die surface roughness on the preform surface quality, we measured the surface roughness of dies and preforms using an optical interferometic profiler (model Contour GT–K1 Optical Profiler Stitching System from Veeco). For the dies, each surface was measured five times at three positions. A magnification of × 5 and measurement area of 1.0 × 1.3 mm was used to determine the roughness. For the preforms, each sample was measured at two different positions close to the end of the preform that was extruded first (the so-called ‘start of extrusion’ end) and with three different magnifications of × 10, × 20 and × 40. To make the measurements with different magnifications comparable the same size of surface area of approximately 120 × 160 µm was used to determine the surface roughness. Before extracting the data for roughness determination, the tilt and the cylindrical shape was removed from the raw data using the profiler software. From these processed optical profiler images, the roughness parameter, Sq (root mean square (RMS) roughness) as defined in  was determined. Note that due to the curved preform surface and the stripes the error in the surface roughness was relatively high (30%).
Figures 3 and 4 show optical profiler images and Sq surface roughness of the surfaces of a 3D-printed Ti alloy die and a conventionally machined stainless steel die. The 3D-printed die exhibited a flat external surface that was media blasted and a flat external surface that was conventionally machined after 3D-printing, which enabled surface roughness measurement for media blasted and machined surfaces of 3D-printed dies. The Sq surface roughness value of 8.2 ± 0.2µm for the media blasted surface agrees with the surface roughness value of 10 µm estimated by the die manufacturer (Section 2). The Sq surface roughness value of 1.5 ± 0.2µm for the machined surface is consisted with the Sq value of 1.9 ± 0.1µm for the conventionally machined stainless steel die surface (Figs. 3 and 4).
As noted in Section 2, all 3D-printed dies were media blasted. Our surface roughness measurements showed the impact of media blasting for an external surface. The impact of media blasting on internal surfaces could not be measured, but the surface roughness is estimated to be between that of a surface out of the 3D printer and that of a media blasted external surface, i.e. between 20 and 10 µm. Thus, the internal surfaces of the 3D-printed dies exhibit approximately an order of magnitude higher surface roughness compared with the machined surfaces of the conventionally machined stainless steel die and the smoothed internal surfaces at the die exit for the 3D-printed Ti alloy die used here.
The high surface roughness of the internal surfaces of the 3D-printed dies of trials A and B resulted in stripes visible at the preform surface with the naked eye. Note that these stripes were not observed for any F2 glass preforms extruded through conventionally machined dies of designs II and III. These stripes were particularly pronounced at the end of the preform that was extruded first (the so-called ‘start of extrusion’ end). Due to gravity, the mass of the preform already extruded causes tensile force on the soft glass emerging from the die exit, which exerts a surface smoothing effect, resulting in a smoother preform surface as the extrusion proceeds. Figures 4 and 5 show optical profiler images and Sq surface roughness values for the preforms close to the ‘start of extrusion’ ends. In agreement with the preform inspection with the naked eye, the two preforms of trials A and B that were extruded through 3D-printed dies with untreated, high-roughness, internal surfaces and that showed the pronounced stripes exhibit high surface roughness of 70 ± 30 nm. Note that the preform surface roughness is approximately 2-3 orders of magnitude lower compared with the internal die surfaces, demonstrating the fire-polishing effect of extrusion due to glass flow at elevated temperatures. The preform of trial D extruded through conventionally machined die exhibits significantly lower surface roughness of 28 ± 4 nm, in agreement with the lower surface roughness of conventionally machined die surfaces compared to 3D-printed, untreated, internal surfaces. The preform of trial C extruded through the 3D-printed die with smooth internal surfaces at the die exit exhibits lowest surface roughness of 8.7 ± 0.2 nm. This result indicates that smoothing of the short-length, internal surfaces at the die exit is sufficient to obtain preform surface quality comparable with preforms extruded through conventionally machined dies.
The die surface roughness measurements infer that the internal surfaces at the die exit are comparable for 3D-printed die C and conventionally machined die D. However, the preform extruded through die C has lower surface roughness than the preform extruded through die D. As we could not measure the internal die exit surfaces for the die C and D, one possible reason for the difference in preform surface roughness could be a difference in die surface roughness for the actual dies used.
As described in Section 3, the extrusion of the two preforms through machined dies required different extrusion force due to different internal die structure and thus friction for the glass flow. Despite the different extrusion force, both preforms exhibit comparable surface roughness of 28 ± 4 nm (Figs. 4 and 5). The surface roughness measurement is in agreement with the observation that no pronounced stripes were observed with the naked eye for preforms extruded through machined dies. The independence of the preform surface roughness on the extrusion force confirms that the different preform surface roughness of trials A,B relative to trials C,D is caused by different die surface roughness at the die exit and not by the different forces used for trials A,B relative to trials C,D.
5. Summary and conclusions
We investigated, for the first time, the use of 3D-printed metal dies for the extrusion of lead-silicate glass preforms. The 3D-printed dies withstood the temperature of 565 °C and forces up to 20 kN, demonstrating that 3D-printed dies have sufficient mechanical strength to enable extrusion of optical preforms. The glass flow through the 3D-printed dies showed the same Poiseuille flow behavior with zero slip coefficient at the glass/die interface as for conventionally machined dies.
The relatively rough surface of the 3D-printed dies prior to treatment of internal surfaces at the die exit resulted in preforms with higher surface roughness compared with the preforms extruded through the machined dies. However, finishing of the easily accessible internal surfaces upstream from the die exit using machine tooling was sufficient to reduce the surface roughness to values similar to that of preforms extruded through machined dies. This result demonstrates that only the finish of the internal die surfaces upstream from the die exit over short length of 1-2mm determines the surface quality of a preform. We attribute this result to the glass viscosity being sufficiently low within the die to enable fast glass flow, which erases the history of roughness introduced earlier in the die. Therefore, in general, finishing of the easily accessible internal die surfaces at the die exit can be applied to improve preform surface roughness for all dies independent of the die fabrication technique. This method of finishing the internal surface at the die exit does not have severe practical limitations as the minimum size of the orifices at the die exit is limited by achieving glass flow at sufficiently low temperatures and force to maintain good optical quality of the glass. Hence the die orifices are usually >0.5 mm for the smallest dimension, enabling surface finishing methods such as machining and polishing.
Despite using different metals (Co-Cr-superalloy, tool steel, Ti alloy, stainless steel), the dies considered in this study demonstrated the same glass flow behavior for F2 glass. Of particular importance is the result for the Ti alloy as this is a widely used metal for 3D printing  but which has not been used for extrusion dies.
The successful use of 3D-printed dies for the extrusion of preforms with a tube structure and the achievement of good preform surface finish by finishing the surfaces at the die exit demonstrates that 3D-printed dies are a viable alternative to conventionally machined dies. The use of 3D-printed dies offers new die designs with arbitrary feed hole dimensions including 3D spirals to tailor the glass flow in different sections of the die. For example, electro-optical fibers  and nonlinear fibers with tailored dispersion  require holes with significantly different sizes at specific locations to each other in the fiber cross-section, which in turn necessitates dies having blockages with significantly different distances to each other and to the die wall to extrude corresponding preforms. To date, conventional machining has limited feed holes to having straight and parallel side walls, resulting in different localized flow rates between the blockages and between blockage and die wall for such dies, which limited the hole size differences and positions that could be achieved in a preform extruded through such a die. In contrast, 3D printing offers complete freedom in the shape of feed holes, allowing the flow in the various regions within a die to be balanced. The flexibility in the 3D geometry of 3D-printed dies also offers dies with arbitrary blockage shapes and directions, opening up preform structures with arbitrarily shaped holes arranged in any arbitrary pattern that could not previously be extruded or even contemplated to date.
We acknowledge the South Australian Government for funding this work under the Premier’s Research and Industry Fund – International Research Grant program. T. Monro acknowledges the support of an ARC Laureate Fellowship. This work was performed in part at the Optofab and Materials nodes of the Australian National Fabrication Facility utilizing Commonwealth and SA State Government funding. We wish to thank the following persons at the University of Adelaide: Herbert Foo and Roman Kostecki for conducting extrusion trials, Wenqi Zhang for measuring surface roughness of one of the preforms, and Piers Lincoln for discussion about 3D printing.
References and links
1. T. Monro and H. Ebendorff-Heidepriem, “Progress in microstructured optical fibers,” Annu. Rev. Mater. Res. 36(1), 467–495 (2006). [CrossRef]
2. H. Ebendorff-Heidepriem and T. M. Monro, “Analysis of glass flow during extrusion of optical fiber preforms,” Opt. Mater. Express 2(3), 304–320 (2012). [CrossRef]
3. J. Bei, T. M. Monro, A. Hemming, and H. Ebendorff-Heidepriem, “Reduction of scattering loss in fluoroindate glass fibers,” Opt. Mater. Express 3(9), 1285–1301 (2013). [CrossRef]
4. G. Tao, S. Shabahang, E.-H. Banaei, J. J. Kaufman, and A. F. Abouraddy, “Multimaterial preform coextrusion for robust chalcogenide optical fibers and tapers,” Opt. Lett. 37(13), 2751–2753 (2012). [CrossRef] [PubMed]
5. H. Ebendorff-Heidepriem, T. Monro, M. A. van Eijkelenborg, and M. C. J. Large, “Extruded high-NA polymer microstructured fiber,” Opt. Commun. 273, 133–137 (2007). [CrossRef]
6. S. Atakaramians, S. Afshar V, H. Ebendorff-Heidepriem, M. Nagel, B. M. Fischer, D. Abbott, and T. M. Monro, “THz porous fibers: design, fabrication and experimental characterization,” Opt. Express 17(16), 14053–15062 (2009). [CrossRef] [PubMed]
7. S. Atakaramians, K. Cook, H. Ebendorff-Heidepriem, S. Afshar V., J. Canning, D. Abbott, and T. M. Monro, “Cleaving of extremely porous polymer fibers,” IEEE Photonics Journal 1(6), 286–292 (2009). [CrossRef]
8. V. V. R. K. Kumar, A. K. George, W. H. Reeves, J. C. Knight, P. St. J. Russell, F. G. Omenetto, and A. J. Taylor, “Extruded soft glass photonic crystal fiber for ultrabroad supercontinuum generation,” Opt. Express 10(25), 1520–1525 (2002). [CrossRef] [PubMed]
9. P. Petropoulos, H. Ebendorff-Heidepriem, V. Finazzi, R. Moore, K. Frampton, D. J. Richardson, and T. M. Monro, “Highly nonlinear and anomalously dispersive lead silicate glass holey fibers,” Opt. Express 11(26), 3568–3573 (2003). [CrossRef] [PubMed]
10. J. Y. Y. Leong, P. Petropoulos, J. V. H. Price, H. Ebendorff-Heidepriem, S. Asimakis, R. C. Moore, K. Frampton, V. Finazzi, X. Feng, T. M. Monro, and D. J. Richardson, “High-nonlinearity dispersion-shifted lead-silicate holey fibers for efficient 1-µm pumped supercontinuum generation,” J. Lightwave Technol. 24(1), 183–190 (2006). [CrossRef]
12. X. Feng, F. Poletti, A. Camerlingo, F. Parmigiani, P. Horak, P. Petropoulos, W. H. Loh, and D. J. Richardson, “Dispersion-shifted all-solid high index-contrast microstructured optical fiber for nonlinear applications at 1.55 microm,” Opt. Express 17(22), 20249–20255 (2009). [CrossRef] [PubMed]
13. W. Q. Zhang, S. Manning, H. Ebendorff-Heidepriem, and T. M. Monro, “Lead silicate microstructured optical fibres for electro-optical applications,” Opt. Express 21(25), 31309–31317 (2013). [CrossRef] [PubMed]
14. H. Ebendorff-Heidepriem, Y. Li, and T. M. Monro, “Reduced loss in extruded soft glass microstructured fibre,” Electron. Lett. 43(24), 1343–1345 (2007). [CrossRef]
15. H. Ebendorff-Heidepriem, S. C. Warren-Smith, and T. M. Monro, “Suspended nanowires: Fabrication, design and characterization of fibers with nanoscale cores,” Opt. Express 17(4), 2646–2657 (2009). [CrossRef] [PubMed]
16. K. J. Rowland, H. Ebendorff-Heidepriem, S. Afshar, and T. M. Monro, “Antiresonance guiding in soft-glass hollow-core microstructured fibres; fabrication and spectra properties,” Australian Conference on Optical Fibre Technology (ACOFT‘2009), Adelaide, 29 Nov – 3 Dec 2009, paper 161.
17. S. C. Warren-Smith, H. Ebendorff-Heidepriem, T.-C. Foo, R. Moore, C. Davis, and T. M. Monro, “Exposed-core microstructured optical fibers for real-time fluorescence sensing,” Opt. Express 17(21), 18533–18542 (2009). [CrossRef] [PubMed]
18. H. T. C. Foo, H. Ebendorff-Heidepriem, C. J. Sumby, and T. M. Monro, “Towards microstructured optical fibre sensors: surface analysis of silanised lead silicate glass,” Journal of Materials Chemistry C 1(41), 6782–6789 (2013). [CrossRef]
19. K. J. Rowland, H. Ebendorff-Heidepriem, S. Afshar, G. Tsiminis, and T. M. Monro, “Extruded soft glass single-ring hollow core fibres”, Australian and New Zealand Conference on Optics and Photonics (ANZCOP’2013), Fremantle, Western Australia, Australia, 8–11 Dec 2013.
20. M. Trabelssi, H. Ebendorff-Heidepriem, K. C. Richardson, T. M. Monro, and P. F. Joseph, “Computational modeling of die swell of extruded glass preforms at high viscosity,” J. Am. Ceram. Soc. (accepted).
22. X. Feng, T. M. Monro, V. Finazzi, R. C. Moore, K. Frampton, P. Petropoulos, and D. J. Richardson, “Extruded singlemode, high-nonlinearity, tellurite glass holey fibre,” Electron. Lett. 41(15), 835–836 (2005). [CrossRef]
23. X. Feng, W. H. Loh, J. C. Flanagan, A. Camerlingo, S. Dasgupta, P. Petropoulos, P. Horak, K. E. Frampton, N. M. White, J. H. V. Price, H. N. Rutt, and D. J. Richardson, “Single-mode tellurite glass holey fiber with extremely large mode area for infrared nonlinear applications,” Opt. Express 16(18), 13651–13656 (2008). [CrossRef] [PubMed]
24. M. R. Oermann, H. Ebendorff-Heidepriem, D. J. Ottaway, D. G. Lancaster, P. J. Veitch, and T. M. Monro, “Extruded microstructured fiber lasers,” IEEE Photon. Technol. Lett. 24(7), 578–580 (2012). [CrossRef]
25. H. Ebendorff-Heidepriem, K. Kuan, M. R. Oermann, K. Knight, and T. M. Monro, “Extruded tellurite glass and fibers with low OH content for mid-infrared applications,” Opt. Mater. Express 2(4), 432–442 (2012). [CrossRef]
26. H. Ebendorff-Heidepriem, P. Petropoulos, S. Asimakis, V. Finazzi, R. C. Moore, K. Frampton, F. Koizumi, D. J. Richardson, and T. M. Monro, “Bismuth glass holey fibers with high nonlinearity,” Opt. Express 12(21), 5082–5087 (2004). [CrossRef] [PubMed]
27. W. Q. Zhang, H. Ebendorff-Heidepriem, T. M. Monro, and S. Afshar V., “Supercontinuum generation in bismuth microstructured optical fiber with three zero dispersion wavelengths,” Opt. Express 19, 21135–21144 (2011).
28. H. T. Munasinghe, A. Winterstein-Beckmann, C. Schiele, D. Manzani, L. Wondraczek, S. Afshar V., T. M. Monro, and H. Ebendorff-Heidepriem, “Lead-germanate glasses and fibers; a practical alternative to tellurite fore nonlinear fiber applications,” Opt. Mater. Express 3, 1488–1503 (2013).
29. E. T. Y. Lee, “Development and characterisation of phosphate glasses for athermalisation,” PhD thesis, University of Southampton (2004).
30. D. Furniss and A. Seddon, “Towards monomode proportioned fibreoptic preforms by extrusion,” J. Non-Cryst. Solids 256-257, 232–236 (1999).
31. S. D. Savage, C. A. Miller, D. Furniss, and A. B. Seddon, “Extrusion of chalcogenide glass preforms and drawing to multimode optical fibers,” J. Non-Cryst. Solids 354(29), 3418–3427 (2008). [CrossRef]
32. Z. G. Lian, Q. Q. Li, D. Furniss, T. M. Benson, and A. B. Seddon, “Solid microstructured chalcogenide glass optical fibers for the near- and mid-infrared spectral regions,” IEEE Photon. Technol. Lett. 21(24), 1804–1806 (2009). [CrossRef]
34. K. Bhowmick, H. P. Morvan, D. Furniss, A. S. Seddon, and T. M. Benson, “Co-extrusion of multilayer glass fiber-optic preforms: Prediction of layer dimensions in the extrudate,” J. Am. Ceram. Soc. 96(1), 118–124 (2013). [CrossRef]
35. K. Itoh, K. Miura, I. Masuda, M. Iwakura, and T. Yamashita, “Low-loss fluorozirco-aluminate glass fiber,” J. Non-Cryst. Solids 167(1-2), 112–116 (1994). [CrossRef]
36. H. Ebendorff-Heidepriem, T.-C. Foo, R. C. Moore, W. Zhang, Y. Li, T. M. Monro, A. Hemming, and D. G. Lancaster, “Fluoride glass microstructured optical fiber with large mode area and mid-infrared transmission,” Opt. Lett. 33(23), 2861–2863 (2008). [CrossRef] [PubMed]
37. H. Ebendorff-Heidepriem, D. G. Lancaster, K. Kuan, R. C. Moore, S. Sarker, and T. M. Monro, “Extruded fluoride fiber for 2.3μm laser application”, International Quantum Electronics Conference (IQEC)/The Conference on Lasers and Electro-Optics (CLEO) Pacific Rim Conference, Sydney, Australia, Aug-Sep 2011.
38. J. Bei, T. M. Monro, A. Hemming, and H. Ebendorff-Heidepriem, “Fabrication of extruded fluoroindate optical fibers,” Opt. Mater. Express 3(3), 318–328 (2013). [CrossRef]
39. C. A. G. Kalnins, H. Ebendorff-Heidepriem, N. A. Spooner, and T. M. Monro, “Radiation dosimetry using optically stimulated luminescence in fluoride phosphate optical fibres,” Opt. Mater. Express 2(1), 62–70 (2012). [CrossRef]
40. C. R. Garcia, R. C. Rumpf, H. H. Tsang, and J. H. Barton, “Effects of extreme surface roughness on 3D printed horn antenna,” Electron. Lett. 49(12), 734–736 (2013). [CrossRef]
41. S. Leuders, M. Thöne, A. Riemer, T. Niendorf, T. Tröster, H. A. Richard, and H. J. Maier, “On the mechanical behaviour of titanium alloy TiAl6V4 manufactured by selective laser melting: Fatigue resistance and crack growth performance,” Int. J. Fatigue 48, 300–307 (2013). [CrossRef]
42. R. Hölker, A. Jäger, N. B. Khalifa, and A. E. Tekkaya, “Controlling heat balance in hot aluminum extrusion by additive manufactured extrusion dies with conformal cooling channels,” International Journal of Precision Engineering and Manufacturing 14(8), 1487–1493 (2013). [CrossRef]