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Synchrotron Radiation (SR) instrumentation poses a special challenge and pushes the envelope for mechanical design of its components. High peak power density of 100 kW/mrad2, high angular collimation (100 μrad) and small source size (˜50 μm), grazing angle geometry and extended arm length (˜10 m) are common to third generation SR sources and create design challenges comparable only to powerful lasers, (Proceedings of MEDSI 2002 conference, Sep. 5-6, 2002, Argonne National Lab, Argonne, Ill., USA). UHV requirements, remote control and sometimes a radiation harsh environment certainly do not simplify such a task.
Soft x-ray beamlines (50-2500 eV) typically have a pre-focusing optical stage, and such mirrors work well as a filter and reduce the power load to downstream components. Monochromatization is performed by grating at grazing incidence, and requires slits for spatial separation of “unwanted radiation”. Entrance/exit directions are fixed and allow the slits to be mounted on a stable platform (K. Kaznacheyev, CLS#6.2.76.4 SM project: Power load and tolerances for optical components; K. Kaznacheyev, I. Blomqvist, E. Hallin, S. Urquhart, D. Loken, T. Tyliszczak, T. Warwick, A. P. Hitchcock; Principles of optical design of the SM beamline at the CLS, proceedings of SRI2003 conference, ed. T. Warwick, AIP 2004). Power density is still kept very high (especially at focusing points) and limits the choice of materials. Precise, backlash free mechanical design with submicron accuracy is required and often calls for a direct reading of blade position.
In addition, blades need to be electrically insulated (to provide a beam monitoring) and placed in a UHV bakeable vessel with out-of-vacuum actuation. Fortunately, tungsten's soft x-ray attenuation coefficient does not exceed 0.4 μm even at high energy, which means that even a thin blade is opaque to soft x-rays and there is no radiation hazard as with hard x-ray applications (K. Kaznacheyev, CLS#6.2.76.15, ID10 chicane magnet layout and the SM performance in quick polarization exchange mode; Petukhov and Hartnett, Advances in Heat Transfer, vol. 6, Academic Press, NY, 1970). Such requirements are quite common and this makes us believe that there is need for a generic design of precise optical slits for soft x-ray beamlines.
UHV version of flexure based, precise optical slit, compatible with requirement s of modern soft x-ray SR beamlines, has been designed by ADC. The following article describes requirements, design principles and mechanical performance of a slit on a basis of an extended thermal and structural FEA analysis.
The invention as described herein with references to subsequent drawings, contains similar reference characters intended to designate like elements throughout the depictions and several views of the depictions. It is understood that in some cases, various aspects and views of the invention may be exaggerated or blown up (enlarged) in order to facilitate a common understanding of the invention and its associated parts.
Provided herein is a detailed description of one embodiment of the invention. It is to be understood, however, that the present invention may be embodied with various dimensions. Therefore, specific details enclosed herein are not to be interpreted as limiting, but rather as a basis for the claims and as a representative basis for teaching one skilled in the art to employ the present invention in virtually any appropriately detailed system, structure, or manner.
SM beamline has no entrance slit, but an exit slit on each of the SM branches: Photo Electron Emission Microscope (PEEM) and Scanning Transmission X-ray Microscope (STXM). These exit slits need to be water cooled, as the maximum power load at zero order is ˜60 W. The exit slits have dual purpose: first to determine the energy window that is delivered to each experiment; second, to serve as the source point for optics which perform further demagnification. Because both the elliptical refocusing mirror of PEEM branch and Zone Plate for STXM branch are stigmatic optics, the displacement along the beam between the vertical and horizontal blade pair shall not exceed the depth of focus of the two downstream optics (5 mm).
The STXM phase acceptance (numerical aperture*exit slit size width) for ultimate spatial resolution should not exceed λ (photon wavelength), and so for a 200 μm diameter Zone Plate at 3 m from the exit slit and 2000 eV photon beam, the exit slit opening will be as small as 6 μm. For the PEEM branch, the e−-source projection to the exit slit will result in a spot size ˜150(h)*7(v) (to 15 depending on grating setting) microns. The nominal resolving power of 3000 this corresponds to slit sizes of 70-20(v) microns, depending on energy and grating settings. A total 0.5 mm total slit opening is required. For the SM beamline, the nominal along-the-beam location of the exit slit does not depend upon the energy and grating setting, but to facilitate the beamline alignment, the exit slit will be placed on an X-Y linear slide and should have a flat bottom.
All beamline components are UHV compatible (˜10−10 torr) and can withstand bake-out to 150° C. for 48 hours, and retain performance characteristics upon cooling to room temperature.
Because of the small range of blade movement and required accuracy of travel, a monolithic flexure 16 design is used, seen in
During assembly alignment, milling machine fixture clamps featuring off centric mounting bolts are used to hold the fine adjustments while the mount block is secured. The blades will have been pre aligned to better that 2 mrad parallel accuracy during installation. A small offset (50 μm) allows a complete close of the blades with small overlapping along beam direction. A 1 mm thickness of ceramic provides electrical resistance between the blade and grounded flexure. This is still thin enough for sufficient thermal conductivity from blade to flexure.
To provide sufficient heat conductance an array of parallel flexure elements 20 are used to connect the slit flexure to the main body. For a given thickness of the monolith, thermal conductivity is proportional to the thickness of the linkage throat, but so is material stress, and force required scales as the cube. Several small linkages are a viable means of providing sufficient cross sectional area for conduction. Balancing requirements for conduction, stress, actuation force, and fabrication costs, we have limited the number of heat-conducting flexure elements to 24 per blade and the thickness of 0.25 μm. The actuation flexure is made thicker, (0.3 μm) and includes a longer throat because of higher stresses and no thermal conduction requirement.
The vertical slit 21, can be seen in
The overall dimensions of the Slit Unit are: upstream to downstream flanges (2¾″ OD CF) is 460 mm, but a more compact design is possible. Slit actuation 29 is made to protrude up and to the right as seen along the beam and cooling water enters from left and exits up. The main slit cube is 194 mm wide×200 mm high×150 mm parallel to the beam and has a flat bottom with tapped holes 31 for mounting on a linear stage. The overall weight is approximately 40 kg.
Available actuators for the slit unit were not found to meet all requirements for this application and so a new actuator design was produced,
This high performance, linear actuator features a precision preloaded ball screw 35, mounted concentric with the non-rotating output and perpendicular to the mounting surface of its rugged housing 33. This configuration makes it equally adept at push-only or push-pull type applications at significantly higher loads than competitive offerings. There are of no reflected loads to the preloaded linear slide that provides guidance. Therefore the accuracy potential of its micro-stepping motor and high precision linear encoder are met throughout its loading envelope. Optical switches (internal) 36 indicate travel limits and zero position. A non-traversing tapered connection attaches to your shaft for pull-push mode, or a steel tooling ball 37 is mounted for push-only mode. A manual knob 38 allows sensitive touch-off zero confirmation in push mode applications.
A number of designs were analyzed in order to optimize actuator loads, flexure stresses, and heat transfer characteristics of the monolithic flexure assembly 16. This design optimization of the monolithic flexure included material selection, several geometric aspects and cooling passage 39 placement within the chamber 26. A graphic representation of the subject design can be seen in
Cooling of the monolithic flexure is accomplished via conduction through the GlidCop flexure structure and the 304 SS vacuum chamber 26 with subsequent convection to the water flow within the cooling passages 39 in the vacuum chamber,
In order to calculate the convective heat transfer coefficient(s) associated with the water flow in the cooling passages, several flow conditions were assumed ranging from laminar to potentially turbulent flow. The water properties, flow conditions and convection coefficients are listed in
where: f=(0.790 ln ReD−1.64)−2. (2)
The convection coefficient is defined as
h=kNu
D
/D
H. (3)
ReD and Pr are defined in
ANSYS 8.0® was used for all finite element simulations. Higher order quadratic elements were employed for each phase of the analysis. For a given level of mesh discretization, quadratic elements typically yield higher accuracy of results when compared to linear elements. Additionally, curved boundaries can be modeled precisely with higher order elements and tetrahedral elements may be used where topographically required without compromising solution accuracy.
Owing to the 2-D nature of the flexure, a plane strain model was utilized for the structural analysis. Approximately 10,250 elements and 33,700 nodes comprised the model. Mesh density in the thin areas of the monolithic flexure was optimized to insure convergence of stress results in these areas as they are subjected to the greatest amount of bending stress by design; ½ symmetry was assumed for the structure. Appropriate displacement boundary conditions (Ux=0) were applied to the vertical symmetry edges on the left side of the model. Full displacement restraints were applied at the mounting holes 40 on the outer periphery of the flexure (5 larger holes). Applied loads consisted of displacements, corresponding to the limits of actuator travel (+/−0.34 mm), and were applied to the lower left vertical edge along with the symmetry condition previously imposed. Linear elastic material properties were assumed for the GlidCop,
The 3D nature of the steady state heat transfer analysis required a corresponding three dimensional finite element model, see
Three flow regimes were evaluated for their cooling effectiveness. Although not shown in the figure, convective heat transfer coefficients were specified on the inner walls of the cooling passages as listed in
Because of the considerable computational resources required (memory in particular) to solve the thermomechanical problem using the 3D flexure model, the 2D model from the structural analysis was utilized for this simulation. The temperature distribution on the center-plane of the 3D thermal model, corresponding to minimal flow conditions, was superimposed on the 2D structural model. This is a reasonable approximation since the temperature variation through the thickness of the flexure was not seen to be significant, less than 1-2° C. The thermomechanical simulation was completely linear in nature, both from a geometric and material perspective.
A graph of the maximum equivalent stress vs. actuator motion can be seen in
As mentioned previously, three flow rates were evaluated for their effect on the cooling of the monolithic flexure. Based on the range of estimated Reynolds numbers for each cooling passage, the flow regimes are referred to as laminar, transitional and turbulent. The results from the heat transfer analysis are summarized in
At a total flow rate of 2.7 gal/min, which has been termed transitional flow, wall temperatures for the cooling passages are typically well below 26° C. validating the assumption of constant properties for the water at this volume rate of flow (and greater). For the minimal flow rate evaluated, the average wall temperature of the cooling passages was seen to be as high as 35° C. Although this is not excessive, its effect on the properties of the water could potentially increase the calculated temperature distribution within the flexure by several degrees. The thermal resistance is seen to be fairly evenly split between the flexure and the bolted interface/vacuum chamber for this flow condition.
The intent of the thermomechanical analysis is to determine the thermally induced deformation of the monolithic flexure at steady state operating conditions. This is an important aspect of the design with respect to slit blade clearance in the fully closed position and potential angular deviation. In the area of the slit blade, the maximum horizontal displacement is −27 μm. This indicates that a slit blade clearance of 54 μm would be required of the design. Further, results show that under these coolant flow conditions, the angular deviation of the slit blade will be 0.19 mrad, well below the +/−2 mrad allowed.