Design and implementation of an integrated and rapidly assembled infrared three-band optical system

1.   INTRODUCTION

In recent years, increasing complexity and variety of detection and imaging targets have led to a surge in demand for multi-band real-time imaging detection systems. Multi-band infrared imaging technology offers enhanced adaptability to intricate environmental conditions and boosts the success rates of detection, tracking, and recognition when compared with single-band infrared imaging systems^[1-3]. Additionally, multi-band infrared imaging systems are capable of capturing multi-dimensional spectral imagery, which allows for the collection of differential information at various wavelengths^[4,5]. This enables the acquisition of valuable information about the target, with spectral band fusion technology further enhancing the quality of imaging detection.

To date, most reported infrared multi-band imaging systems are limited to dual-band configurations^[6-9], which have difficulty in capturing the full spectrum of radiation characteristics across the short-wave, mid-wave, and long-wave spectral ranges. Multi-band optical systems commonly employ dispersive elements, such as prisms and dichroic mirrors, for light separation^[10,11], but these elements suffer from coating efficiency limitations that result in the loss of incident light energy, substantially reducing the energy received by the photosensitive surface and necessitating a larger optical aperture to ensure adequate information gathering. Consequently, multi-band composite optical systems tend to be bulky and complex, which hampers their widespread adoption.

In response to the need for rapid deployment, short development cycles, and cost-effectiveness in the aviation, space, and astronomical sectors, metal-based mirrors, particularly those made from aluminum alloy, have emerged as a promising solution due to their high performance-to-cost ratio, and have been integrated into space launch payloads^[12-13].

With the advancement of third-generation infrared mercury cadmium telluride (HgCdTe) detector technology, the efficiency of multi-band optical systems in capturing light has significantly improved. There is an urgent requirement to design a lightweight, rapidly deployable, and versatile multi-band composite imaging detection system that caters to these new third-generation infrared detectors.

This paper presents a multi-band common optical path design that aligns with the imaging mechanisms of novel detectors, aiming to achieve an ultra-compact infrared three-band system tailored for the characteristics of short-wave, mid-wave, and long-wave infrared targets commonly used in space detection imaging. Using the efficient manufacturing capabilities of single-point diamond turning and adopting an integrated optical-mechanical design, we have developed a rapidly assembled three-band infrared optical system. It features a simple, compact structure, and holds significant potential for application in the aviation, space, and astronomical detection fields, which are characterized by a demand for miniaturization and mass production.

2.   THREE-BAND OPTICAL SYSTEM DESIGN

2.1   Three-Band Imaging Working Principle

The operation of the three-band composite imaging optical system is shown in Fig. 1. Third-generation infrared HgCdTe detectors are capable of multi-band layered imaging, with readout circuitry capable of distinguishing imaging information across various wavebands^[14-16]. These detectors can disperse and color-separate the incident light energy. The optical composite imaging of the novel detection system requires just a single optical dispersion to achieve simultaneous imaging across all three bands. This streamlines the design of our optical system, enabling us to concentrate on optimizing imaging quality across a broad spectral range, and on the appropriate matching and selection of optical materials. The three-band infrared detector adopts a common Dewar structure, in which one band is imaged separately through beam splitting, and the other two bands achieve co-focused imaging using a planar dual-color stacked detector. This detector comprises photo-sensitive elements absorbing different infrared bands, arranged or stitched apart on the same plane. The advantage of this type of detector is that devices for different bands can be designed and manufactured integrally in the same process.

Figure  1.  Schematic showing the operation of a three-band composite optical system.

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2.2   Construction of the Initial Optical System Framework

The optical system design presented in this work encompasses three operational wavelength bands: Short-wave, mid-wave, and long-wave infrared. The requirement for these bands to share a common focal plane poses a significant design challenge. Additionally, the broad operational wavelength range restricts the choice of optical materials, with only a select few (such as germanium, zinc sulfide, and certain chalcogenide glass materials) being suitable for use.

The design process commences with the preliminary selection of a catadioptric optical system configuration. We use a two-mirror system with the R-C (Ritchey-Chrétien) configuration, favored for its ability to correct both primary spherical aberration and coma simultaneously. The structure is shown in Fig. 2, where it is compared with the traditional Cassegrain configuration, offering a more streamlined approach to establishing the initial system design. The focal point F₁ of the primary mirror is located at the back of the secondary mirror. The part blocked by the secondary mirror is represented by a red dotted line. In the two-mirror system, the red incident light rays are reflected from the primary mirror to the secondary mirror and then converge at the focal point F₂ of the system. The semi-heights of the primary mirror and the secondary mirror are h₁ and h₂ respectively. The distance between the primary mirror and the secondary mirror is d. The distances from the vertex of the secondary mirror to the two focal points are l₂ and l'₂ respectively, and the focal length of the primary mirror is denoted as f'₁. For the rear group, we select materials that meet the requirements of three-band imaging, while the front and rear groups are matched and jointly optimized^[17].

Figure  2.  Schematic diagram of initial optical design parameters for the R-C two-mirror system.

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On completion of the initial structure for the front group, we expand the imaging field to accommodate short, mid, and long-wave infrared requirements. This is achieved by using a combination of chalcogenide glass, multispectral ZnS, and Ge materials to achieve comprehensive and balanced correction of aberration across the three bands. Furthermore, in consideration of tolerances and manufacturability, the addition of high-order aspheric surfaces can enhance aberration correction capabilities, resulting in a more compact optical-mechanical structure.

For a telephoto system with an infinite object distance, the optical parameters are defined as follows:

Object distance: ${l}_{1}=\infty ,{u}_{1}=0$;

Primary-secondary mirror interval: $d$;

Half-diameter of the primary mirror: ${h}_{1}$;

Half-diameter of the secondary mirror: ${h}_{2}$;

Blockage ratio: $\alpha ={h}_{2}/{h}_{1}$ ;

Secondary mirror magnification: $\beta ={l}^{'}_{2}{{}}/{l}_{2}$;

Primary mirror conic coefficient: ${e}_{1}^{2}$ ;

Secondary mirror conic coefficient: ${e}_{2}^{2}$ ;

Seidel spherical aberration coefficient: S_Ⅰ;

Seidel coma coefficient: S_Ⅱ;

Seidel astigmatism coefficient: S_Ⅲ;

Seidel field curvature coefficient: S_IV;

Seidel distortion coefficient: S_V;

The radius of the primary mirror: R₁;

The radius of the secondary mirror radius:

For a two-mirror system, only five monochromatic aberrations need to be considered. The third-order aberration coefficients corresponding to spherical aberration, coma, astigmatism, field curvature, and distortion are as

For a two-mirror system, where R₂=2α/(β+1), ${n}_{1}^{\prime}={n}_{2}=-{1}$, let $h_{1}={1}$,$f_1=1$ and $\theta=-1$, we can derive the following relationships: $f_1=1/\beta$, $u_1=u_2=\beta$, $u'_2=1\ .$

From these relationships, we can obtain the optical system parameters, The initial parameters for two mirrors are detailed in Table 1.

Table  1.  Initial parameters solution for two mirrors

Parameters	Mirror1	Mirror2
Pi	$- \dfrac{{{\beta ^3}}}{4}$	$\dfrac{{{{\left( {1 - \beta } \right)}^2}\left( {1 + \beta } \right)}}{4}$
Wi	$\dfrac{{{\beta ^2}}}{2}$	$\dfrac{{1 - {\beta ^2}}}{2}$
$\prod_{{{i}}}^{ }$	$\beta$	$- \left( {1 + \beta } \right)$
${\phi _i}$	$- \beta$	$\dfrac{{1 + \beta }}{\alpha }$
Ki	$- \dfrac{{{e_1}^2}}{4}{\beta ^3}$	$- \dfrac{{{e_1}^2{{\left( {1 + \beta } \right)}^3}}}{{4{\alpha ^3}}}$

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By substituting the parameters from the above equations into the Equation (1)−(5), we can derive

If S_I=S_II= 0, we then obtain

2.3   Optical System Design Specification Parameters

In response to the specific requirements of a three-band infrared detection system, we present an example design, employing a detector with a 256 × 256 cryogenic focal plane array and pixels measuring 30 µm× 30 µm. The design specifications of this system are detailed in Table 2. The imaging quality of the optical system is characterized by a modulation transfer function (MTF).

Table  2.  Parameters of the example optical system

Parameters	Specifications
Wavelength Bands	Short-wave: 2.5–2.9 µm Mid-wave: 3.7–4.8 µm Long-wave: 7.7–9.5 µm
F#	2
Field of view	3°×3°
Focal length	146.5 mm
Optical system length	200 mm
Weight	410g
Image quality	Short-wave: MTF ≥ 0.40 @ 17 lp/mm Mid-wave: MTF ≥ 0.35 @ 17 lp/mm Long-wave: MTF ≥ 0.30 @ 17 lp/mm

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2.4   Design Results of the Optical System

The optical system employs a secondary imaging structure to achieve cold shield matching, with the intermediate image plane located between the primary and secondary mirrors. To minimize the impact of radiation from the mirror itself on detection performance, cold shield matching is a critical design consideration. The aperture stop is positioned at the cold shield of the detector, and the entrance pupil is controlled to be as close as possible to the position of the primary mirror, thereby reducing the effective diameter of the primary mirror. To enhance the optimization of higher-order aberrations, the primary and secondary mirror surfaces are modified from quadratic surfaces to higher-order aspheric surfaces.

Considering the design specifications, system obscuration requirements, and overall length envelope, the initial optical system structure is optimized. Due to the impact of central obscuration on optical transfer function, the surface of the mirrors is controlled to have an obscuration ratio better than 1∶4 under the premise of meeting the system image quality requirements.

We use the CODE V software package to optimize the initial structure. The completed optical system is shown in Fig. 3. The system consists of a two-mirror system (primary mirror and secondary mirror), a lens correction group, and a detector assembly. The incident light beam is reflected by the primary and secondary mirrors and then passes through the lens group to form an image on the focal plane of the detector. The detector is divided into two photosensitive surfaces. The reflected light acts on the photosensitive surface FPA₁, while the transmitted light acts on the photosensitive surfaces FPA₂ and FPA₃. The combination of lens focal lengths is positive-negative-positive-negative-positive, with materials in a sequence of IG24-ZnS-IG24-Ge-IG24. Lens materials can be quickly machined using the single point diamond turning method.

Figure  3.  2D layout optical path of the confocal plane for the three-band infrared optical system.

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The surface obscuration in this system is optimized to 22.6% following design refinement. The total length of the optical system, measured from the secondary mirror to the detector window, is meticulously controlled to be 104.5 mm, with the tube length ratio maintained around 0.7, creating an ultra-compact design for the three-band system.

We evaluate the imaging capabilities of the optical system at the Nyquist frequency (17 lp/mm) across the three wavelength bands of short-wave infrared (2.5–2.9 µm), mid-wave infrared (3.7–4.8 µm), and long-wave infrared (7.7–9.5 µm), as shown in Fig. 4. Considering the MTF variation of the whole field of view, we set a total of 5 fields of view from C₁ to C₅ for evaluation. For the short-wave infrared range of 2.5–2.9 µm, the edge field MTF exceeds 0.65 at the Nyquist frequency. In the mid-wave infrared range of 3.7–4.8 µm, the edge field MTF is greater than 0.58 at the Nyquist frequency. For the long-wave infrared range of 7.7–9.5 µm, the edge field MTF is better than the Nyquist frequency by 0.3. The MTF of the optical system is nearly at the diffraction limit, with a 100% efficiency in cold shield matching, ensuring excellent imaging quality.

Figure  4.  Three-Band optical MTF curves, with the x-axis showing spatial frequency and the y-axis showing normalized MTF. C₁– C₅ correspond to the MTF curves of different fields of view respectively.

(A) Long-wave MTF curves. (B) Mid-wave MTF curves. (C) Short-wave MTF curves.

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2.5   Tolerance Allocation Results

Rapid assembly of optical systems can reduce manufacturing costs, and tight tolerances need to be considered during the design phase. To accommodate rapid assembly, it is crucial to validate whether the optical system possesses a reasonable tolerance margin. This makes it essential to estimate the tolerance impact domain of the optical system and identify sensitive tolerance items.

With the inevitable presence of manufacturing and alignment errors in production processes, the manufacturing errors of the three-band infrared optical system include tolerances in the surface figure error, radius error, and quadratic coefficient error. Alignment errors include decentration error, tilt error, and spacing error of mirrors and lenses. Since the optical system is a coaxial rotationally symmetric system, only radial decentration and axial spacing displacement need to be considered. During the optical system alignment, the primary mirror (P-M) is fixed as a reference, so it only needs to be analyzed for manufacturing errors during tolerance analysis. Alignment tolerances need to be considered for the secondary mirror (S-M) and lenses.

This is accomplished through the use of Monte Carlo analysis in the CODE V software, which is employed to analyze the tolerance. Subsequently, the tolerances are adjusted to reasonable ranges based on engineering feasibility, and they are then redistributed and evaluated^[18].

The rapid manufacturing and quick assembly characteristics of the three-band infrared optical system are crucial objectives. While meeting the image quality requirements for infrared three-band imaging, it is necessary to reduce the tolerance sensitivity of the optical system, reducing the dependence on assembly accuracy and facilitating rapid assembly. Tolerances are readjusted and allocated from both manufacturing and assembly perspectives to ensure that the optical system has good feasibility in manufacturing and assembly processes. The preset values of the tolerances are listed in Table 3.

Table  3.  The tolerance assignment results of the optical system, index is the tolerance item, P-M is the primary mirror, S-M is the secondary mirror, and Lens is the lens group

Index	P-M	S-M	Lens
Figure error	0.07λ (λ=632.8 nm)	0.07λ (λ=632.8 nm)	0.07λ (λ=632.8 nm)
Radius error	0.02 mm	0.02 mm	0.01 mm
Quadratic coefficient error	0.01	0.01	None
Spacing error	Reference	0.02 mm	0.02 mm
Decenter	Reference	0.015 mm	0.015 mm
Tilt	Reference	1′	40″

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After setting the tolerance distribution, we assess the MTF across diverse fields of view for the short-wave, mid-wave, and long-wave infrared spectral bands, accounting for the established tolerances. Within the bounds of these predefined tolerances, the short-wave band can maintain an MTF of at least 0.37 with a 97.7% probability. For the mid-wave band, the MTF is expected to be above 0.45 with a 97.7% probability, while the long-wave band is projected to achieve an MTF of over 0.38 with a 97.7% probability. The tolerance MTF curve is shown in Fig. 5.

Figure  5.  Monte Carlo tolerance analysis results for the three-band infrared optical system after adding optical tolerances. C₁–C₅ correspond to the wavefront differential tolerance analysis curves of different fields of view respectively.

(A) Probability curve showing long-wave MTF reduction. (B) Probability curve showing mid-wave MTF reduction. (C) Probability curve showing short-wave MTF reduction.

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3.   RAPIDLY ASSEMBLABLE OPTO-MECHANICAL STRUCTURE DESIGN

Given the small tube length ratio of the optical system, an ultra-compact structural layout is necessary to achieve the lightweight and miniaturizable opto-mechanical structure. With the growing application and advancement of metal optics, the trend toward integrated opto-mechanical design is increasingly demonstrating its benefits. First, metal optics can be rapidly processed using the single-point diamond turning (SPDT) method, enabling the fabrication of more complex aspheric surfaces. Second, the backplate support structure is directly integrated into the metal mirror, with pre-designed connection holes.

This approach facilitates the joint design of the mirror body and the support structure, eliminating the need for additional screws for tightening or adhesive layers for bonding. Moreover, the use of the same material for both the mirror body and the support structure eliminates the issue of mismatched thermal expansion coefficients, facilitating the realization of an athermal optical system. This simplifies the complexity of the opto-mechanical design while enhancing the environmental adaptability of the components^[19].

3.1   Overall Configuration of the Structure

The thin catadioptric optical system poses specific challenges in the spatial arrangement of the mirrors and structural components. Ultra-compact layout and integrated design are effective strategies for achieving lightweight and miniaturized design^[20,21].

The optical system in this study is a coaxial catadioptric configuration, featuring a very short distance between the primary mirror and the rear lens group. Consequently, the primary mirror support structure and the lens group are arranged in a radially spaced staggered layout. The optical surface of the primary mirror and the structural components are designed with an integrated approach. This approach integrates the secondary mirror support with the secondary mirror tube, the primary and secondary mirrors with the support backplate, and the primary flange with the lens tube. Our design reduces the number of components from the 10 used in traditional designs to 6, improving the efficiency of system assembly, while minimizing assembly errors and degrees of freedom. The basic composition of the opto-mechanical structure is depicted in Fig. 6.

Figure  6.  Opto-mechanical structure of catadioptric optical system.

(A) Schematic diagram of components designed for conventional structures. (B) Schematic diagrams of components designed for our newly proposed structures that can be rapidly assembled.

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3.2   Integrated Mirror Structure Design

The design must ensure that the mirror body possesses high rigidity to prevent deformation under static and dynamic loads. Additionally, it is essential to design the support structure of the mirror to ensure thermal and mechanical stress isolation between the mirror surface and the external mounting interface. To enhance the adaptability of the mirror assembly to external environments, a flexible hinge structure is required to dissipate thermal and mechanical loads. For this optical system, which is intended for use in aerospace environments with significant thermal and mechanical loads, it is crucial to maintain the stability of the mirror position. This can be achieved using a high-torque mounting mechanism.

To ensure the rapid assembly of the mirror, we propose an integrated mirror design structure. The structural design focuses on three specific aspects. Threaded holes on the mirror body should be avoided. These can lead to issues such as uneven thread straightness, thread gaps, and machining errors, which can directly affect the final mounting surface quality and introduce installation stresses. Maintaining adequate clearance between the mounting surface and the mirror is also important. If space is limited along the axis, the use of a flexible hinge slot should be considered to increase the force transmission path and reduce sensitivity to changes in mirror surface quality. Finally, a flexible structure can accommodate installation torque. This flexible mirror structure should be able to accommodate the installation torque in the direction tangential to the screw, ensuring that installation stresses are released during the mounting process.

In line with these principles, we have designed a flexible hinge structure capable of accommodating large installation torques. This structure integrates the mirror body with the flexible hinge, allowing the hinge to deform along the tangential direction of the screw preload to dissipate installation stresses. A three-dimensional model of the mirror, which is constructed from aluminum alloy, is shown in Fig. 7.

Figure  7.  3D model of the primary mirror structure design, with three-dimensional axes shown by XC, YC, and ZC.

(A) Front view of the mirror. (B) Back view of the mirror.

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3.3   Integrated and Optimized Design of Support Structures

Given the structural layout and weight constraints of the optical system, an optimized design of the secondary mirror support tube is essential. One commonly employed method for achieving this is material topology optimization, which involves the strategic removal of non-load-bearing materials. The Solid Isotropic Material with Penalization (SIMP) model, a variable-density approach for solid isotropic materials with penalties, is a prevalent optimization technique in continuous structure topology optimization. This method employs a stiffness constraint to minimize the volume of the structure during optimization^[22,23].

The optimization of the support structure is a static optimization problem, in which the objective function aims to minimize the structural compliance (or maximize stiffness by minimizing strain energy), with the structural volume ratio serving as the constraint function. The mathematical model for this^[24] is given by Equations (13) and (14).

where C is the compliance of the mirror, with the objective function of minimizing strain energy. K is the global stiffness matrix, U is the global displacement vector, F is the global load vector, F^T is the F matrix operation, U^T is the transposed U matrix.

where s.t. is the optimized constraints, ρ is the unit density, V is the design domain volume constraint, $f$ is the volume fraction, and V₀ is the design domain volume.

We use the optimization module of the HyperMesh software package (Altair, Troy, Michigan, USA) to complete the topology optimization analysis under given boundary constraints. In light of the ultra-compact layout of the optical system and the topology optimization results, the topology-optimized secondary mirror support is modeled for engineering implementation, resulting in the final secondary mirror support structure shown in Fig. 8.

Figure  8.  Support design model based on topology optimization.

(A) Density contour map of the topology optimization. Blue shows material removal and red shows material retention. The red ring near the base is outside the designed area. (B) 3D model of the lens barrel remodeled with reference to the optimization results.

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4.   ANALYSIS OF THE FEASIBILITY OF RAPID MANUFACTURING AND ASSEMBLY OF LIGHTWEIGHT MIRRORS

Metal mirrors with apertures of 200 mm or less can achieve surface flatness with an accuracy better than 1/10λ (where λ = 632.8 nm) by using single-point diamond turning (SPDT) processing. To further improve surface precision, it is essential to carefully consider the forces acting on the mirror during the manufacturing process. A primary mirror with a significant thickness ratio (i.e., ratio of diameter to thickness) must be evaluated for its manufacturability and ease of assembly, as these factors are critical for creating a three-band optical system.

During the SPDT process, the surface quality of the mirror is primarily influenced by centrifugal deformation of the mirror surface and flatness deformation from the mirror mounting. The forces affecting the mirror during processing are typically dominated by centrifugal forces and the cutting force of the single-point tooling, which can range from approximately 0.3 to 1.5 N^[25]. For mirrors with rotationally symmetric structures and adequate support rigidity, any processing stresses can be considered negligible.

4.1   Analysis of Machining Centrifugal Forces

Given the typical operating conditions of single-point diamond turning, with a spindle speed of 500 revolutions per minute (RPM), we conduct an analysis of centrifugal deformation during the single-point diamond turning process of the primary mirror. By removing the rigid body displacement and angle, surface shape fitting is carried out using the MATLAB software package (MathWorks, Natick, Massachusetts, USA), and results are shown in Fig. 9.

Figure  9.  Centrifugal force analysis result of SPDT.

(A) Displacement contour diagram showing centrifugal deformation. (B) Mirror deformation contour of the primary mirror after removing the displacement and rotation of the rigid body. Red shows large displacement, and blue represents downward displacement.

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We find that the flexibility of the primary mirror is particularly well-suited for rotational turning on a single-point diamond lathe. Despite the flexibility of the structure, the centrifugal deformation generated by the diamond turning process is minimal. The root mean square (RMS) deformation of the primary mirror is 0.66 nm, indicating excellent rapid manufacturing processability.

4.2   Rapid Assembly Flatness Adaptability Analysis

In addition to the processing deformation, the final surface quality of the mirror must also account for the impact of installation^[26]. We analyze the flatness adaptability of the primary mirror under typical installation conditions, using a three-point mounting configuration with positions A, B, and C. The simulated installation conditions are divided into three cases:

Case 1: Installation point A has a flatness of 3 µm, B has a flatness of 0 µm, and C has a flatness of 0 µm.

Case 2: Installation point A has a flatness of 3 µm, B has a flatness of 1 µm, and C has a flatness of 0 µm.

Case 3: Installation point A has a flatness of 3 µm, B has a flatness of 2 µm, and C has a flatness of 0 µm.

The final analysis results are presented in Fig.10. The primary mirror has good adaptability to an installation with a flatness of 3 µm, satisfying the installation requirements of the infrared optical system.

Figure  10.  Cloud maps of RMS deformation on the primary mirror surface after removing the displacement and rotation of the rigid body. Red shows large deformation, and blue shows small deformation.

These cloud maps show deformation underr (A) the case 1 condition, (B) the case 2 condition, and (C) the case 3 condition.

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Under different installation conditions, the installation deformation of the mirror may vary, and the best adaptability is to the installation of a single point that is either raised or depressed. For rapid assembly, it is necessary to consider the yield rate of assembly, so the worst deformation result for mirror installation adaptability must be determined based on assembly tolerance requirements. The least favorable situation is equivalent to the three-point installation condition, in which one installation point is used as a reference zero point with the other two installation points having high and low deviation, respectively. This causes the installation surface to form a wave-like deformation. This working condition, as the most extreme, is considered more reliable for analysis results. It is also evident that controlling the planar consistency of two installation points can effectively reduce the deformation of the installed mirror surface.

5.   RAPID MANUFACTURING AND ASSEMBLY ADAPTABILITY VERIFICATION

5.1   Rapid Manufacturing of the Mirror

The primary and secondary mirrors of the main system are fabricated using single-point diamond turning (SPDT) processing. During the manufacturing process, the flatness of the mounting surface is meticulously controlled, and an auxiliary support method is employed to further reduce the impact of centrifugal deformation on the mirror surface. As shown in Fig. 11, the surface flatness of the turned mirrors is better than 1/20λ (where λ = 632.8 nm). The mounting surface and the mirror surface are machined in the same clamp, ensuring that the mirror surface tilt perfectly matches the back mounting. This unifies the processing and alignment reference, reducing the complexity of subsequent optical system alignment. During optical alignment, only centration and decentration adjustments are necessary, eliminating the need to adjust the mirror tilt, which significantly simplifies the alignment process.

Figure  11.  Processing, measurement process, and measurement results of the primary mirror. Red shows large deformation, and blue shows small deformation.

(A) The completed primary mirror. (B) Test results of the surface shape of the primary mirror, including RMS and Peak and valley (PV) profile errors, power errors, and sample sizes.

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5.2   Rapid Installation Adaptability Verification

To verify the simulation analysis conclusions from Section 4.2, we select the primary mirror for rapid installation plane adaptability validation and match it with the three installation conditions mentioned in the simulation analysis. This verifies the adaptability of the primary mirror under different planarity conditions, with three-point installation. The installation surface of the primary mirror is ground and treated to simulate various planarity conditions that may occur during the installation process. Interferometric surface profile tests are conducted on the primary mirror under installation conditions 1, 2, and 3, which are likely to be encountered during mirror installation.

The surface accuracies of the primary mirror (shown in Fig. 12) under the three installation conditions all meet the specified tolerance requirement of 0.07λ (λ = 632.8 nm). Under installation condition 1, the change in the primary mirror surface is 0.012λ (λ = 632.8 nm), equivalent to 7.6 nm; under installation condition 2, the change is 0.025λ (λ = 632.8 nm), equivalent to 15.8 nm; and under installation condition 3, the change is 0.019λ (λ = 632.8 nm), equivalent to 12.0 nm. Measurement discrepancies arise from the challenges in artificially creating installation planarity defects that perfectly match the simulated conditions. However, the consistency of the installation planarity change patterns can be assessed from the trends observed. The test results confirm that the primary mirror has good surface adaptability under the installation condition of 3-micron planarity, while also validating that the flexible design of the primary mirror is more adaptable to the most demanding installation environments, such as wave-like surface deformation. During the process of mirror installation plane machining, controlling the consistent planarity of at least two installation points can effectively reduce the deformation of the mirror surface after installation, providing a definite reference for engineering, manufacturing, and implementation.

Figure  12.  Flatness adaptability verification results. Red shows large deformation, and blue shows small deformation for the primary mirror.

(A) case 1 conditions, (B) case 2 conditions, and (C) case 3 conditions.

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6.   OPTICAL SYSTEM PERFORMANCE TESTING

The optical lens is precision-assembled on the collimator, resulting in a three-band infrared system with a total weight of just 408 g. The compact volume of the system measures only Φ112 mm × 117 mm. Following the alignment and adjustment process, we conduct a comprehensive performance test of the entire system. The MTF of the system is measured using a specialized transfer function measuring instrument, with the test carried out using a mid-wave infrared detector.

Detailed test results are shown in Fig. 13. The tangential direction MTF is 0.50 @ 17 lp/mm, and the sagittal direction MTF is 0.405 @ 17 lp/mm. The discrepancy is caused by one lens having a surface accuracy that exceeds the tolerance, while still meeting the requirement that the mid-wave MTF should be better than 0.4. Due to project schedule constraints, the surface error of the lens was not corrected or adjusted. We find that the optical system meets the stringent requirements for high-quality three-band imaging.

Figure  13.  Medium wave infrared MTF test. The red curve shows the meridian MTF curve, and the blue curve shows the sagittal MTF curve. Inset are sample points of the transfer function curve.

(A) Photograph of the optical system being tested. (B) MTF curves of the measured optical system by software screenshots.

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7.   CONCLUSIONS

To meet the imaging detection requirements for short-wave, mid-wave, and long-wave infrared spectra, we present an ultra-compact and lightweight three-band optical system. Based on third-generation detector imaging technology, it enables simultaneous co-focused imaging across all three infrared bands.

Initially, based on the imaging principles of the third-generation multi-color detectors, we develop a design methodology for a three-band co-focused optical system. Starting with a two-mirror system with chalcogenide glass optical materials, we achieve a wide-tolerance optical system. To enhance assembly efficiency, we introduce an integrated design approach for optical and structural components, using the rapid manufacturing characteristics of metal mirrors and topological optimization to minimize the number of assembly parts and degrees of freedom in the optical system. We confirm the rapid assembly features of the mirrors through simulations and tests. The total weight of the assembled optical system is less than 408 g, with a compact volume of only Φ112 mm × 117 mm. The full field-of-view transfer function in the mid-wave infrared band exceeds 0.405 @ 17 lp/mm, satisfying the stringent requisite imaging criteria.

Our system offers distinct advantages over traditional multi-band infrared optical systems, such as a compact and lightweight design, a broad tolerance range, and the capability for quick assembly. The design and fabrication of this optical system serve as a valuable reference for the advancement of lightweight and miniaturized multi-band infrared optical systems in aerospace and astronomical exploration applications.