2.1.

Subscale Beryllium Mirror Demonstrator

The SBMD project produced a 0.53-m diameter beryllium mirror with a 20-m radius of curvature (ROC) mounted on a solid Be support structure built by Ball Aerospace (BATC) (see Fig. 1).¹⁶ It was cryogenically tested multiple times at MSFC and provided invaluable experience and learning.¹⁷ For example, SBMD had cryogenic quilting (cryo-quilting), but the mechanical model did not predict any cryo-quilting. After several design iterations, how to properly model the cryo-quilting was learned.¹⁸ Using this knowledge, new rules were defined for how to design lightweight beryllium mirrors without cryo-quilting. These new design rules were successfully proven on AMSD. Additionally, SBMD taught valuable lessons on how to design cryogenic interfaces that do not distort the mirror surface shape. SBMD was also used to certify that the Webb gold coating, uncorrectable surface figure error, and creep were at TRL-6 (see Table 7). However, great caution is advised whenever extrapolating technical performance results from small mirrors to large mirrors. Given its size and design, SBMD was significantly stiffer than either AMSD or the Webb flight mirrors.

Fig. 1

SBMD.

SBMD was also important as the first use of O-30 beryllium for a cryogenic mirror. In 1996, the state-of-the-art cryogenic mirror was the 85-cm Spitzer Telescope made of I-70 beryllium¹⁹ with a diffraction limited performance of $5\mu \mathrm{m}$. But I-70 Be was not a good choice for the NGST primary mirror. Because it was produced using a mechanical pulverization process, its powder had irregular grain shapes. This irregularity limited how densely the powder could be packed into a hot isostatic pressure (HIP) can—which limited the maximum size mirror that could be made. Also the irregular grain shapes resulted in large CTE inhomogeneity. The solution was O-30 Be developed by Brush Wellman for the Air Force in the late 1980s. Because O-30 Be is a spherical powder material, it has a high packing density (thus allowing for hot isostatic pressuring of larger billets), and its CTE distribution is very uniform (which results in smaller cryo-distortion and higher cryo-stability). Also because O-30 Be has a lower oxide content than I-70 Be, it can achieve a smoother polished surface (i.e., less scatter). The ability to HIP a meter class billet was demonstrated in the late 1990s via the VLT secondary mirror. By 1999, Brush-Wellman had full production capability sufficient for the NGST program.

This author’s personal lessons learned from SBMD include the following.

2.3.

NGST Mirror System Demonstrator

The NMSD project was the most technically aggressive study. It sought to explore the limits of light weighting and successfully showed what does not work. NMSD clearly demonstrated the important roles that CTE and mechanical stiffness play in the ability to design and manufacture a stable mirror system. NGST produced two 1.6-m hexagonal shaped spherical mirrors with a 20 m ROC and areal density $<15\mathrm{kg}/{\mathrm{m}}^2$. The two mirrors were manufactured by Composite Optics Inc. (COI) and the University of Arizona. The COI mirror was a thin glass sheet bonded to a rigid graphite composite structure. The Arizona mirror was a thin glass sheet attached to a graphite composite structure via 166 actuators.

Both NMSD mirrors took significantly longer to make and achieved significantly lower cryo-performance than expected. The causes for these results were assessed to be CTE mismatch and inhomogeneity; too low of areal density (i.e., too low stiffness); and overly complex designs.

This author’s personal lessons learned from NMSD include the following.

2.4.

Advanced Mirror System Demonstrator

The AMSD study was designed to explore the most likely NGST mirror technologies at an appropriate scale. Its success formed a basis for estimating Webb ambient and cryogenic performance, manufacturability, schedule, cost, and risk. Given the importance of large lightweight mirrors to many government missions, AMSD was a joint NASA and Department of Defense program. Although some mission requirements were divergent, the pooling of resources provided greater funding to explore the technology landscape more widely and deeply.

A critical element of the AMSD program was competition. Competition between ideas and vendors resulted in a rapid TRL advance of modern, large-aperture lightweight cryogenic space mirrors. AMSD followed a phased down-select approach. Phase 1 awarded contracts to five different vendors to study and develop designs for a total of eight different mirror architectures. The best four were funded for fabrication in phase 2. Ball Aerospace, Goodrich, and Kodak were the winning vendors.^9,20–25 All of these mirrors were 1.3 to 1.4-m point to point—just the size needed to produce a segmented primary mirror 6 to 8 m in diameter—and had an areal density of $\sim 15\mathrm{kg}/{\mathrm{m}}^2$.

BATC, building upon their earlier SBMD work, developed an open-back beryllium mirror that incorporated ROC and mirror position control utilizing flight-like cryogenic actuators (see Fig. 2).^20–23 A key element of the BATC approach is that the mirrors required cryo-null polishing to remove cool down cryodistortions.

Fig. 2

Beryllium AMSD mirror.

Goodrich proposed two high-authority mirror concepts consisting of a face sheet (one concept was shallow-ribbed glass and the other was silicon carbide) supported on an array of displacement actuators. The displacement actuators would be used to correct for cool down distortion.²⁴ The face sheets would be fabricated via stress polishing on a mandrel. Early in its design phase, because of projected cost and schedule overruns, NASA terminated the SiC concept.

Kodak (now L3-Harris) fabricated a semirigid mirror system that utilized a closed-back all-glass cellular-core mirror along with a few force actuators to correct low-order mirror distortions that occur during cool down to cryogenic temperatures (see Fig. 3).^9,25,26 The Kodak approach also assumed cryo-null polishing to remove both correctable and uncorrectable cryo-deformations.

Fig. 3

ULE^® glass AMSD mirror.

After Northrop Grumman (NGC) was selected as the Webb prime contractor in 2002, the Goodrich effort was terminated due to incompatibility with the NGC Webb architecture. This allowed the remaining funds to be focused on the BATC and Kodak mirrors. Both mirrors were successfully cryo-tested, and their cryogenic performance was characterized. Findings of the cryo-testing included the following: a properly designed beryllium mirror substrate will not have cryo-quilting; a properly designed mirror mount will not introduce low-order cryo-deformation; and cryo-deformation is the result of CTE nonuniformity in the mirror substrate.

This author’s personal lessons learned from AMSD phase 2 include the following.

It is the assessment of this author that AMSD (and the broader NGST mirror technology development effort) was successful because of the following.

3.1.

PhaseCAM

Testing long ROC mirrors and overcoming the limitations of atmospheric turbulence and mechanical motion had long been an interest of this author. One solution to this problem was phase-measuring interferometry (PMI) at $10.6\mu \mathrm{m}$.²⁷ The infrared wavelength was insensitive to temperature induced index variations in the atmosphere and small amplitude mechanical vibrations. But infrared interferometers lack visible wavelength sensitivity. Another solution (developed at Breault Research Organization) was high-speed PMI with a 340-Hz Reticon CCD camera.²⁸ If the atmospheric turbulence and mechanical motion are sufficiently slow, they can be sampled and removed via averaging—as long as the sampling interval was longer than the atmosphere’s random-walk correlation time.

In 1987, Keck contracted with BRO to use this high-speed PMI system to test their primary mirror segments at ITEK. Because of their mirror’s long focal length and off-axis near-parabolic optical prescription, the total optical test air-path for each segment was 48 m and required five reflective bounces—twice off mirror segment and three times off the autocollimation flat (ACF). This air path was too long for ITEK’s LUPI (laser unequal path interferometer) or a commercial Fizeau temporal phase-shifting interferometer, and there was too much mechanical motion. The high-speed camera system helped, but the ultimate solution was a pseudo common-path test setup that this author had seen in Norm Cole’s optical shop. The mechanical motions were mitigated by rotating the LUPI reference flat 45 deg to send the reference beam along the same path as the test beam (bouncing off the ACF and a flat attached to the segment test stand). Atmospheric turbulence was mitigated by averaging 64 measurements taken on a multiminute cadence to avoid atmospheric turbulence “random-walk” correlation.²⁸ In the process, another problem was discovered, a problem that would be critical for absolutely characterizing the LIGO reference flats and for testing the Webb segments. Because the CCD camera frame capture was not instantaneous, laser frequency drift introduced a phase shift error from the start of the frame readout to the end.

Now for serendipity, before joining NASA, while visiting MetroLaser on another matter, this author saw on a table in a back lab a breadboard setup of a “real-time” interferometer producing a phase map of a flame plume. Its potential was immediately obvious, and Bernie Seery, Pre-Phase-A Study Manager, agreed to fund a risk reduction experiment. Newly incorporated 4D Vision Technology was given a $60 K contract to build an interferometer to NASA’s specification. They delivered the first ever PhaseCAM (Fig 5) in just 6 months and it worked great. Its resolution was $512\times 512$; its repeatability was 1.2 nm rms; and over a 20-m air path, its measurement uncertainty was 5 nm rms. It was put into use immediately, and Webb could not have been made without the 4D PhaseCAM technology.

Fig. 5

4D PhaseCAM #1.

3.2.

Absolute Distance Meter

Regarding measuring the ROC, a common technique is to use a distance measuring interferometer on a lens bench to measure the radius. But this technique would not work for the NGST mirrors because it requires a displacement measurement from “cats-eye” to the center of curvature. During a NASA fact-finding trip to SSG Walthan MA, a Leica theodolite was observed with an interesting distance measuring technology, but its accuracy was only $\pm 5\mathrm{mm}$, and the NGST specification was $\pm 0.1\mathrm{mm}$. So a development effort with Leica was funded and resulted in the absolute distance meter (ADM). The ADM was used to measure and set the ROC on all development and flight mirrors. (And maybe its funding led to Leica’s commercial DISTO handheld laser distance measuring devices.)

5.1.

Performance Subcommittee

Vendor primary mirror design concepts were evaluated for their impact on the Webb OTE level 2 performance requirements. Technical performance criteria were divided into two categories: mandatory and secondary. Mandatory criteria were defined as mirror performance factors that influence the ability of the OTE to meet Webb level 2 requirements. Secondary criteria were other factors that influenced OTE performance but were not directly traceable to level 2 requirements. Each mirror concept was characterized as to its ability to exceed, meet, or not meet mandatory evaluation criteria (see Table 2).

Table 2

Mandatory performance criteria.

The Performance Subcommittee assessed that meeting key Webb Level 2 requirements would be very challenging but that the beryllium mirror provided significant performance advantages for Webb.

Regarding the science mission timeline, both mirrors were assumed to be able to satisfy the 10-year requirement. The only difference between them was that, if the end-of-life temperature of the OTE is 4 K different than the beginning of life temperature, then the ULE^® mirror’s wavefront error (WFE) would degrade $\sim 8\mathrm{nm}$ rms. Regarding launch survival, it was assumed that both mirrors could be designed for the required vibroacoustic environment.

Regarding the collecting area, Webb requires each mirror segment to be polished to within 5 mm of the physical aperture. Both mirror teams had some difficulty with this on AMSD, but the ULE^® team had slightly more difficulty.

Compliance with the optical performance parameters was derived based on AMSD ambient and cryogenic test results. To facilitate a head-to-head comparison, a specific protocol was defined to ensure that all data were analyzed identically. This protocol defined in the data flow how to correct for CGH distortion; remove low-order aberrations; compensate for gravity sag; mask, clip and threshold the data; remove test setup induced misalignment aberration; and how much real cryo-deformation aberration would be allowed to be removed based on a limited ability to compensate for such errors by moving mirror segments on-orbit via a hexapod mount.

Based on AMSD results, neither mirror team predicted an OTE WFE of $<117\mathrm{nm}$ rms. The Be team predicted 118 nm rms and the ULE^® team predicted 119 nm rms. However, both vendor teams subsequently figured their AMSD mirrors to a surface quality sufficient to meet this requirement.

A fundamental difference between the Be and ULE^® mirrors were their cryogenic distortion, i.e., mirror shape change from ambient to 30 K. AMSD showed that the ULE^® mirror experienced a larger cryo-distortion than the Be mirror. The total ambient to cryogenic figure change for the Be mirror was 171 nm rms and for the ULE^® mirror was 398 nm rms. As shown in Fig. 6, after performing a simulated hexapod adjustment to remove alignment aberrations, the Be mirror change was 77 nm rms, and the ULE^® mirror change was 188 nm rms. After removing 36 Zernike coefficients, the high spatial frequency residual error is $26\pm 2\mathrm{nm}$ rms for Be and $47\pm 9\mathrm{nm}$ rms for ULE^®. The ULE^® mirror exhibited obvious print through from its core structure. Although it is possible to remove cryo-deformation via cryo-null figuring, the larger magnitude of the ULE^® deformation, as well as its print through, posed more of a risk than that of the Be mirror. Also because this deformation was only recently identified as a result of the AMSD project, there was additional risk that the magnitude and sign of this cryogenic deformation could vary from mirror to mirror.

Fig. 6

AMSD cryogenic distortion results (cryogenic figure–ambient figure).

Another significant difference between the two mirrors was their thermal stability and sensitivity to a thermal operating set point. This sensitivity impacts encircled energy (EE) stability and on-orbit performance as a function of time. Over the anticipated operational temperature range of 30 to 55 K, the Be AMSD mirror experienced a 7 nm rms total surface figure change, whereas the ULE^® mirror’s figure changed by 40 nm rms. Note that the Be change was at or below the measurement sensitivity. As shown in Fig. 7, after removing alignment aberrations, the ULE^® change dropped to 21 nm rms. After removing the first 36 Zernikes, the figure error for Be was 1.6 nm rms and for ULE^® was 4.6 nm rms.

Fig. 7

AMSD cryogenic stability over the thermal operating range (30 to 55 K).

A related risk was the change in the ROC as a function of temperature. The AMSD Be mirror predicted a radius change of $-13\mathrm{mm}$, and a change of $-13.06\mathrm{mm}$ was measured. Its radius change sensitivity over the operating range is only $-0.1\mathrm{ppm}/\mathrm{K}$. The ULE^® mirror predicted a radius change of $+1.4\mathrm{mm}$ and measured a change of $-4.3\mathrm{mm}$. Over the operating range, the radius sensitivity was about $-1\mathrm{ppm}/\mathrm{K}$. Consequently, it was concluded that a ULE^® is more susceptible to uncertainty in operating temperature and thermal gradients than a Be mirror.

Using the measured AMSD optical performance as the basis, the Performance Subcommittee derived the OTE optical performance as measured by the Strehl ratio, point spread function (PSF), EE, and EE stability. Both mirrors predict an excellent Strehl ratio with a substantial margin (Be > 93.9% and ULE^® > 92.8%). The predicted EE for Be was $>74\%$, and the predicted EE for ULE^® was $>72\%$. However, because of the thermal stability of Be, it has a better EE stability over the entire thermal operating range and potential thermal gradient conditions. This was important because EE stability was a financial bonus performance incentive parameter in the prime contract.

Secondary evaluation criteria for each mirror concept were characterized as to its risk (high, moderate, or low) of impacting OTE performance (see Table 3).

Table 3

Secondary performance criteria.

Regarding design issues, all Beryllium design features for Webb (except for the segment size) were equal to or lower risk than what was demonstrated on AMSD. The Be Webb design had fewer pockets than AMSD with thicker ribs, a thicker face sheet, larger corner radii, and larger fillet radii. The ULE^® mirror had several parameters in its design that were higher risk than what was demonstrated on AMSD. In addition to a larger segment size, it would have a thinner front and back plate thickness, larger core depth, and a thinner edge ring. Mass was explicitly excluded as a selection factor, but at the time of the evaluation process, the Be mirror design was 32 kg over budget, and ULE^® would have been 176 kg over budget.

Finally, a new process was proposed for inspecting raw ULE^® glass before it was fabricated into mirrors. The purpose of this process was to mitigate the mechanism thought to explain the thermal deformation effect observed on AMSD. The subcommittee found both the mechanism and new process credible but unproven.

The Performance Subcommittee recommended Beryllium for Webb based on several factors. It was expected to meet all Webb level 2 requirements. AMSD successfully demonstrated most of the critical technology issues needed to scale up to Webb, and Webb design improvements would make the segments more producible with lower risk. The anticipated cryogenic deformation was within the range of what could be cryo-null figured. Beryllium’s excellent thermal properties provided a stable mirror performance over the entire Webb operating temperature range. The subcommittee’s findings on ULE^® were that its AMSD cryogenic behavior was not predictable, the mechanism for that behavior was unproven, and its thermal sensitivity could pose a risk to on-orbit optical performance. However, it was also the assessment of the subcommittee that these findings apply only to cryogenic operation and that all AMSD data support the fact that ULE^® is an excellent mirror material for ambient applications.

5.2.

Manufacturing Process and Test Plan Subcommittee

Each vendor provided detailed schedules for an EDU, 18 flight segments, and 2 spare blanks with 100s to 1000s of elements. Each schedule included critical resource allocations and basis of estimates. AMSD was assumed as the basis for all operations, and detailed traceability matrices were provided to justify all Webb processes. Vendors were instructed to provide detailed justification for any process durations different from AMSD.

From the schedules, the subcommittee selected four critical milestones for assessment: EDU vibe test, EDU completion, first segment, and last segment. EDU structural and acoustic testing were selected to ensure that this critical milestone is accomplished before the nonadvocate review (NAR). EDU completion was a key selection for ensuring that all production steps were fully demonstrated before they were needed on flight hardware. The first primary mirror segment was selected because it was required for the OTE Pathfinder risk reduction activity. The last segment was selected because it defined the critical path to launch.

The manufacturing process and test plans were assessed for adequacy in terms of detail and thoroughness of understanding and adequacy of required equipment. The schedules were assessed for contingency, robustness, and flexibility. Slack to completion was evaluated for each of the four identified milestones. Available workarounds were considered, and a critical path chain analysis was performed. Each schedule was assessed for credibility and risk. The adequacy of justification for the differences between Webb and AMSD was assessed. Risk and mitigation plans were evaluated using Webb project criteria/guidelines.

Using each vendor’s schedule, a comparative probabilistic schedule risk assessment was performed using Risk+© software. The assessment was based on schedule pessimistic and optimistic durations. Pessimistic assessment was based on interrogating the traceability matrix. Optimistic assessment was based on the vendor’s identified process improvements. The assessment indicated Beryllium had 7% to 15% more schedule risk than ULE^® (see Fig. 8).

Fig. 8

(a) Optimistic and (b) pessimistic comparative probabilistic schedule risk for Webb.

The manufacturing process and test plan subcommittee had three findings. First, Beryllium had fewer new or modified steps from AMSD in its proposal than ULE^®. Second, the ULE^® schedule had more slack in the identified program milestones, and its processing flow was appreciably more immune to large perturbations without impacting the Webb critical path. Third, both vendor’s schedules were optimistic and represented risk to the program. The probabilistic risk assessment indicated that Beryllium represented a 7% to 15% greater schedule risk than ULE^®.

5.3.

Facility and Staffing

The facility and staffing subcommittee assessed that there were no major challenges for the ULE^® team and that the biggest challenge for the Be team would be setting up the polishing plant. The biggest equipment challenge for the Be team would be getting the first CNC machine on-line, whereas the biggest challenge for the ULE^® team would be getting small tool machines on-line.

5.4.

Cost Subcommittee

Both vendors provided detailed cost proposals. After adjusting both proposals based on an in-depth review of their basis of estimates, the cost subcommittee found little difference in either proposal’s risk or content. Both vendors proposed a recurring cost that was 35% lower than one might predict using a simple AMSD extrapolation. Additionally, both vendors proposed substantial internal investment. It was the assessment of the subcommittee that much if not all of the cost of the AMSD mirror technology development program was recovered by these cost savings. Finally, based on the relatively greater schedule risk of 7% to 15% for Be, it was assessed that Be had a larger risk of cost overrun than ULE^®. However, this risk was offset by the fact that 45% of the Be proposal was a firm fixed price.

5.5.

Science Subcommittee

The science subcommittee assessed that Beryllium would be the highest performing, lowest technical risk solution. It had superior cryogenic CTE and thermal conductivity—thus providing significant optical performance margin in the event of thermal gradients and bulk temperature set point uncertainty. It also was more forgiving of differences between the thermal conditions for on-orbit and ground testing. The Beryllium mirror would provide, with margin, a PSF that meets both the EE and EE stability requirements.

7.1.

Requirements Flowdown

Technology requirements for the PMSA were derived from level 1 science requirements through level 2 mission requirements and level 3 observatory requirements (see Table 5). Level 1 science requirements were defined in the Webb Program Plan.³² Level 2 mission requirements were defined in the Webb Mission Requirements Document.³³ Level 3 observatory requirements and specific mirror technology component requirements were derived during the phase 2 NGST Observatory Contract and refined after the Prime Contractor (and Implementation Team) was selected. Complete PMSA requirements are defined in the Equipment Specification for Webb PMSA.³⁴

Table 5

PMSA requirement traceability.

A comparison of these PMSA requirements with the pre-Webb state-of-the-art for space telescopes, as defined by Hubble and Spitzer (see Table 6), clearly showed that they were truly well beyond the state-of-the-art. Thus these capabilities were TRL-6 technologies that needed to be demonstrated.

Table 6

Webb mirror technology versus state-of-the-art.

Although there were literally 100’s of engineering specifications necessary to manufacture a Webb PMSA, only a select few were considered technology requiring demonstration: gold coating cryo-survivability, figure thermal stability, areal density, figure launch distortion, primary mirror optical area, surface figure error (including ROC, hexapod, creep and polishing error), and cryogenic performance. Note that this list is not in a priority order, but in the order of their flow down from the level 1 science requirements developed in Tables 5 and 6. The balance of this section details the system engineering logic of how each mirror technology requirement flows from its originating level 1 science requirement.

Although the observatory operating temperature was listed as a key technology, it was really an existence principle. It is the one requirement that pervades all other requirements. To achieve the level 1 science requirement of providing a thermal environment that permits the science instruments to have zodiacal light background limited imaging performance over the wavelength range from 1.7 to $10\mu \mathrm{m}$, the observatory must limit its thermal emissions by operating at a cryogenic temperature of $<50\mathrm{K}$. This directly drives the need to place the telescope at L2, which requires an EELV launch vehicle that demands low areal density mirror segments. This requirement also directly drove all operational thermal requirements, including performance, survival, and stability. Thermal modeling indicated that some of PMSAs might be as cold as 28 K.

Gold coating cryogenic survivability was a relatively minor TRL-6 technology. Level 1 science requirements specified a spectral range of 0.6 to $27\mu \mathrm{m}$. This, in combination with sensitivity, flowed into a level 2 optical transmission requirement, which directly flowed into a PMSA reflectivity requirement. Uncoated polished Beryllium cannot achieve the required reflectivity over the required spectral range. Overcoats of gold, silver, and aluminum were considered. Gold was the best candidate material. It provides excellent reflectivity in the near- and mid-infrared and acceptable performance in the visible. Silver does provide better performance in the visible, but it requires a protective layer to avoid oxidation problems. Aluminum, although common for ground-based visible telescopes, does not have acceptable infrared performance. Gold is a common coating material and thus was not itself a TRL-6 technology. But the cryogenic survival of a gold coating applied to a large O-30 Beryllium mirror had never been demonstrated.

The PMSA surface figure thermal stability was possibly the most important TRL-6 technology and was a key factor in selecting Beryllium as the primary mirror material.^11,33 Level 1 specified that science observations must be able to occur at any position in the celestial sphere. This placed a stability requirement on the EE as the observatory slews, which in practice was a constraint on how much the PSF shape could change due to thermal gradients introduced into the telescope as a function of angle to the sun. At the PMSA level, EE thermal stability is directly determined by the thermal stability of the surface figure shape. Although dozens of engineering issues can contribute to this stability (such as material CTE uniformity and structural design, including actuator athermalization bracket design and bimetallic effects), it was the system level PMSA performance that is the TRL-6 technology. Thus a specific PMSA design implementation must be demonstrated to have cryogenic figure stability of $<0.3\mathrm{nm}$ rms per K, which manifested itself as a maximum surface figure change of 7.5 nm rms from 30 to 55 K.

PMSA areal density was one of the two key technologies identified as requiring significant development effort. The level 1 science requirement of operating the observatory at L2 flowed down to a level 2 requirement that the observatory must be launched via a heavy lift rocket (such as an Areianespace Ariane 5). This placed a mass constraint of 6159 kg on the observatory. The original primary mirror allocation of this mass was 1000 kg, and given that the original telescope collecting area was to be $50{\mathrm{m}}^2$, this placed an areal density requirement on the primary mirror of $20\mathrm{kg}/{\mathrm{m}}^2$. To provide margin, a technology goal of $15\mathrm{kg}/{\mathrm{m}}^2$ was defined. It was this goal that drove the entire mirror technology development program. As the observatory architecture evolved and mass maturity of different observatory elements improved, the PMSA areal density specification was raised to $26.5\mathrm{kg}/{\mathrm{m}}^2$.

The PMSA diameter was the second key technology identified as requiring significant development effort. Originally, an 8 m class primary mirror was required to achieve the desired observatory sensitivity. Given that the observatory needed to operate at L2, that the only way to get to L2 was to be launched on a heavy lift rocket, and that the maximum available shroud diameter was only 4.5 m, it was clear that a segmented and deployed architecture was required. Competing design solutions required segments with diameters ranging from 1 to 3 m. Ground-based observatories (Keck, Hobby-Eberly) and test beds (LAMP, ALOT, and AOSD) had demonstrated the ability to produce segmented telescopes, but their areal densities were too high (70 to $2000{\mathrm{kg}/\mathrm{m}}^2$). Thus a primary focus of the mirror technology development effort was on how to manufacture 1 to 3 m class mirror systems with the required areal density. A key task was to design and demonstrate a substrate that could be manufactured, safely handled, optically finished including ground testing, and integrated into a system that would survive launch—all with an areal density $<20\mathrm{kg}/{\mathrm{m}}^2$. A second issue was the ability to manufacture the substrate blank. Pre-Webb, all large mirrors were glass, which, although acceptable for ambient operation, were less than ideal for a cryogenic telescope. The largest cryo-mirror was the ITTT 0.85-m I-70 Be mirror. Hence, the AMSD program was tasked with demonstrating the ability to manufacture a 1.5-m class O-30 Beryllium mirror blank—as well as the entire mirror system. The Webb PMSA diameter of 1.5-m, which is slightly larger than what was demonstrated on AMSD, was derived from a combination of the level 1 science requirement to have a minimum of $25{\mathrm{m}}^2$ of unobscured optical collecting area and the choice of an 18 segment architecture.

PMSA cryogenic surface figure, creep, launch distortion, and adjustability requirements were derived from performance metrics directly traceable to the level 1 science requirement that the observatory shall be diffraction limited at $2\mu \mathrm{m}$. To achieve the level 1 requirement, the telescope was required to have a residual WFE of $<131\mathrm{nm}$ rms after fixing correctable on-orbit figure errors. To “fix” correctable errors, each PMSA has the ability (at temperatures $<50\mathrm{K}$) to change its ROC and adjust its rigid body position. Detailed error budgeting by Ball Aerospace partitioned the residual WFE between multiple sources, including uncorrectable residual PMSA surface figure error; errors in the ability to adjust all PMSA’s to a common ROC; errors in the ability to phase all PMSA’s into a common primary mirror by correcting PMSA rigid body errors; creep of a PMSA figure as a function of time; and figure change experienced by a PMSA as a function of the launch environment. The result of this process was that each PMSA had to be able to adjust its ROC with a resolution of $\le 10\mathrm{nm}$ PV sag, and each PMSA had to be able to adjust is piston position with a resolution of $\le 10\mathrm{nm}$. Uncorrectable PMSA cryogenic surface figure error—i.e., errors that cannot be corrected by ROC adjustment or sliding the PMSA in the “parent” space with the hexapod—had to be $\le 23.7\mathrm{nm}$ rms at delivery to the OTE integration and test (I&T) process. Also, from the time that a PMSA is delivered for I&T through its end of life, the uncorrected surface figure error from material creep had to be $\le 1.8\mathrm{nm}$ rms. Furthermore, the PMSA uncorrectable surface figure distortion due to launch had to be $\le 2.9\mathrm{nm}$ rms.

An interesting detail of these requirements was the role of material stress/strain and precision elastic limit (PEL) on PMSA design and their connection with figure creep, launch deformation, surface figure error, and areal density. To meet the creep and launch figure change requirements, it was critical that the PMSA substrate had sufficient stiffness to avoid introducing excessive stress/strain into the mirror during optical fabrication. It is the release of this stress/strain from the mirror with time or exposure to the launch environment (vibration and acoustic) that causes undesired figure change. PMSA stiffness is also important for in-process optical testing and I&T; a mirror must have sufficiently small gravity sag that it can be accurately measured in one-g (i.e., on the Earth) while being manufactured for optimized performance in zero-g. So although AMSD demonstrated that an areal density $<20\mathrm{kg}/{\mathrm{m}}^2$ was achievable, a specification of $26.5\mathrm{kg}/{\mathrm{m}}^2$ was necessary to produce a PMSA with sufficient stiffness to meet the other requirements.

7.2.

Requirements Verification

To certify compliance with TRL-6, specific success criteria were established for each critical technology and then confirmed by test. Table 7 lists the critical requirements, their success criteria, and how each was confirmed by test.

Table 7

Mirror technology success criteria.

All PMSA technologies necessary to meet the Webb level 1 requirements were demonstrated to be at TRL-6 via a piecewise methodology. As desirable and recommendable as it might be, TRL-6 was not demonstrated on a single-mirror system. Rather, SBMD, AMSD, flight mirrors, and test coupons were used to mature specific technologies and demonstrate their performance in a relevant environment. For example, SBMD demonstrated gold coating performance at 28 K and cryo-null figuring. Although AMSD was designed to explore fabrication limits associated with areal density and size, it could not certify everything. AMSD did produce a complete mirror system (with a design that is traceable to flight) and tested its performance in a relevant environment from 30 to 50 K. But AMSD was not designed to (nor was it ever intended to) meet Webb launch loads. To survive launch, Webb flight PMSAs were redesigned to have significantly more areal density than AMSD (which made a Webb PMSA easier to fabricate). Additionally, Webb’s flight PMSA design was modified based on lessons learned from AMSD. Therefore, it was necessary to use an actual Webb flight PMSA for vibration and acoustic testing.

The PMSA-110 and PMSA-530 ability of a gold coating to survive the 28 K requirement was verified with SBMD. TRL-6 was demonstrated by performance testing at 30 K and survival testing to 28 K a gold coating deposited on the SBMD mirror. Because cryogenic adhesion of gold on O-30 beryllium was the ability being tested and not the ability to deposit gold coatings on to large mirrors, it was determined that repeating the test with a gold coated AMSD mirror was unnecessary. The deposited gold coating introduced no discernible cryogenic surface figure distortion into SBMD. The uncoated SBMD’s 30 K surface figure was 52.8 nm rms, and its coated 30 K surface figure was 53.9 nm rms³⁵ (Fig. 9).

Fig. 9

Results of SBMD BATC-IRAD coating demonstration.

The PMSA-530 requirement that a PMSA could operate over a 28 to 50 K temperature range was verified with the AMSD mirror system and the Webb flight actuators. TRL-6 was demonstrated by testing the AMSD beryllium mirror system multiple times over operational temperatures from 28 to 50 K to characterize its cryogenic performance. Cryogenic figure stability was characterized. The cryogenic figure and ROC change were demonstrated, and the cryogenic ROC adjustability was demonstrated. TRL-6 was further demonstrated by testing the cryogenic performance of the Webb flight actuators.

The PMSA-170 requirement that a PMSA maintains a surface figure stability of $<0.3\mathrm{nm}$ rms for a 1 K temperature change (7.5 nm rms over a 30 to 55 K thermal range) was verified with AMSD. TRL-6 was demonstrated by measuring the surface shape of the AMSD beryllium mirror system as a function of temperature. The cryogenic surface figure was measured at multiple temperatures and was found to change linearly with temperature. The total surface figure change from 30 to 55 K was 7.0 nm rms or 0.28 nm rms per 1 K temperature change^11,28 (Fig. 10).

Fig. 10

Measured cryogenic figure stability from 30 to 55 K, 7 nm rms (0.28 nm rms/K).

PMSA-410 and PMSA-70 derived requirement that a PMSA could be manufactured with an areal density $<26.5\mathrm{kg}/{\mathrm{m}}^2$ was verified with AMSD and confirmed with Webb flight segments. TRL-6 was demonstrated by calculating the areal density of the AMSD beryllium mirror and an assembled PMSA (Fig. 11) from measurements of their respective masses and physical dimensions. The achieved areal density for the PMSA was $25.8\mathrm{kg}/{\mathrm{m}}^2$. AMSD actually demonstrated the feasibility of manufacturing a mirror system with an areal density of $15.6\mathrm{kg}/{\mathrm{m}}^2$. This was achieved by CNC machining a beryllium mirror substrate with exceptionally thin ribs and facesheet while controlling the introduction of residual stress. Residual stress is very important. It can adversely affect the ability to polish a beryllium mirror to the required surface figure and keep that shape because of long-term figure creep. The higher PMSA areal density requirement (allowed by design maturity, incorporating lessons learned from AMSD and validated with improved modeling) improves manufacturability and reduces risk.

Fig. 11

Assembled (a) AMSD and (b) Webb PMSA mirror systems.

The PMSA-70 requirement that a PMSA could be manufactured with a polished surface area of larger than $1.46{\mathrm{m}}^2$ was verified via a combination of SBMD, AMSD, and EDU. TRL-6 was verified by three specific demonstrations of fact. First, the Webb flight program successfully manufacturing and machined a 1.315 m flat to flat beryllium substrate. Although this may seem trivial now, before the mirror technology development program, there was great uncertainty as to whether or not the manufacture of beryllium substrates of that size was even feasible. Second, AMSD demonstrated the ability to fabricate a 1.2 m flat to flat polished beryllium mirror with a mechanical design and aspheric prescription traceable to Webb. Until it was surpassed by Webb, AMSD was the largest diameter beryllium mirror ever fabricated. Third, SBMD demonstrated the ability to use small tool polishing on a lightweight mirror substrate to within 5 mm of a straight edge.

The PMSA-150 requirement that a PMSA could be polished with an uncorrectable surface figure error of $<23.7\mathrm{nm}$ rms was verified with SBMD and AMSD. TRL-6 was confirmed by verifying two key abilities: (1) the ability to polish a large-aperture low-areal-density aspheric O-30 beryllium mirror to the required specification and (2) the ability to cryo-null figure an O-30 beryllium mirror to have the required figure specification at temperatures $<50\mathrm{K}$. The ability to polish a meter-class highly aspheric lightweight O-30 beryllium mirror was demonstrated on AMSD. AMSD was polished to have an uncorrectable surface figure error of 19.2 nm rms over 97.1% of its aperture (Fig. 12). Achieving a $<20\mathrm{nm}$ rms surface figure was actually the last major task of the AMSD program, and its accomplishment represented a never before demonstrated capability for meter-class lightweight beryllium mirrors. Furthermore, because AMSD had a 10 m ROC, it was a more difficult prescription to polish than Webb segments with their 16 m ROC. The $<20\mathrm{nm}$ rms uncorrectable surface figure was achieved via a small tool computer controlled optical surfacing (CCOS) technology at Tinsley Laboratories in Richmond, CA. Critical to this accomplishment was highly spatial sampled data and precision fiducial registration knowledge. The ability to cryo-null figure such a mirror to yield the required surface figure error at cryogenic temperatures was demonstrated on SBMD. SBMD exhibited a cryo-deformation of $\sim 90\mathrm{nm}$ rms. This shape change consisted of a low-order mount induced error and a high-order quilting error associated with the substrate rib structure. After two cryo-cycles proved that the deformation was stable and repeatable, i.e., that the O-30 beryllium mirror had no apparent creep induced figure change associated with residual stress in the mirror, SBMD was cryo-null figured. The predicted final cryogenic surface figure was 14.4 nm rms. The actual final cryogenic surface error was 18.8 nm rms³⁵ (Fig. 13).

Fig. 12

Final AMSD ambient surface figure is 19.2 nm rms (97.1% of mirror area).

Fig. 13

SBMD predicted cryo-figure of 14.4 nm rms versus actual cryo-null figured cryogenic surface error of 18.8 nm.

Based upon the SBMD success of cryo-null figuring via small tool CCOS technology of both low-order mount induced as well as high-order rib structure quilting, it was determined unnecessary to cryo-null figure AMSD.

The PMSA-370 requirement that a PMSA could be positioned in space with 6 degrees of freedom (DOF) with $<10\mathrm{nm}$ step resolution was verified with AMSD and PMSA components. TRL-6 was demonstrated by test of the PMSA actuator performance at 30 K and analysis of PMSA hexapod motion at 30 K. The Webb cryogenic hexapod mechanism, with its six cryogenic actuators, controlled the 6 DOF position of a mirror segment relative to the Webb telescope backing structure. A seventh actuator was used to deflect the center of the mirror, changing the ROC for that segment. Although the use of a hexapod was not new technology, the actuator step size resolution required at cryogenic temperature was. To meet the hexapod motion resolution and accuracy requirements, the Webb actuators had to be independently capable of $<10\mathrm{nm}$ step size resolution at $<50\mathrm{K}$. This level of motion resolution was achieved when the Webb actuators were operated in their “fine” mode. Webb actuators are dual stage with coarse and fine operating modes. The Webb actuators were developed by BATC, initially under IRAD funding, and then via AMSD, to meet specific mass, stiffness, and performance requirements. These actuators were used for both PMSA hexapod and ROC adjustments.

The key component of the actuator is a cryogenic capable geared stepper motor, which was derived from the gear motor flown on the Spitzer Space Telescope and operated at 4.5 K. TRL-6 capability was demonstrated by characterizing the cryogenic performance from 25 to 35 K of over 24 actuators: 2 actuators via Ball IRAD, 4 actuators via AMSD,³⁶ and 18 Webb engineering unit actuators. All actuators met the resolution requirement with the Webb engineering unit actuators showing a resolution of 7 nm (Fig. 14). Extensive testing of the actuators through a variety of fine-stage step increments verified that the actuator performs single steps, without backlash, to an accuracy of 0.6 nm rms. Finally, flight actuators were installed into a flight hexapod system and exercised at ambient temperature to show basic functionality.

Fig. 14

Actuator single step resolution at 25 K.

The PMSA-1560 requirement that a PMSA cryogenic ROC sag could be adjusted by $<10\mathrm{nm}$ peak-to-valley (pv) was verified with AMSD. TRL-6 was demonstrated by test, analysis, and corollary. PMSA mirrors were designed to adjust their ROC at cryogenic temperatures by expanding or contracting a linear actuator. The actuator, attached to the back center of the mirror, reacts its force via six struts that attach to each mirror corner through a flexured joint. A similar design was implemented on AMSD except that the actuator reacted its force against spreader bars (Fig. 11). Once at 30 K, the AMSD actuator was commanded to execute “coarse-steps” until an ROC sag change was detected. A move of 35 coarse steps resulted in an ROC sag change of 38 nm pv. By analysis, a single AMSD “coarse-step” should result in a sag change of $\sim 1.1\mathrm{nm}$ pv. And a single “fine-step” motion (which is 4.5 times smaller than a coarse step) should result in an ROC sag change of $\sim 0.24\mathrm{nm}$ pv.²⁸ Because of the difference between where the actuator force reacts against the mirror substrate, the distance between those reaction points, and the intrinsic stiffness between AMSD and Webb, a Webb PMSA experiences an ROC sag change that is $\sim 110\%$ larger per linear motion than AMSD experiences. Thus the minimum Webb coarse step is $\sim 1.2\mathrm{nm}$ pv, and the minimum fine step is $\sim 0.27\mathrm{nm}$ pv.

The PMSA-195 requirement that a PMSA could be designed such that its surface figure changes by $<1.8\mathrm{nm}$ rms because of creep was verified with SBMD, AMSD, and Webb flight segments. TRL-6 was demonstrated by test and analysis. Funded via AMSD, Draper Laboratory measured the creep properties of O-30 beryllium.³⁷ Significant creep was measured for samples stressed to 4 and 6 ksi. Neglible creep was measured for samples stressed to 2 ksi or below (Fig. 15). Analysis indicates that 2 ksi of stress will creep 1.8 parts per million over 10 years at room temperature.³⁸ Further analysis indicates that a PMSA with a surface stress of 2 ksi will see a total figure change of $<1.8\mathrm{nm}$ rms during its room temperature life prior to launch and that no figure change due to creep is expected on orbit at cryogenic temperatures. A rule was established that all beryllium components of the Webb PMSA must be designed, processed, and handled in such a way that no component had a residual stress of $>2\mathrm{ksi}$. Additionally, extensive tests were performed under AMSD III to quantify exactly how much stress was introduced into a Be mirror during the machining process at AXSYS and grinding/polishing process at Tinsley. These processes were controlled to limit the residual stress in the final mirrors to $<2\mathrm{ksi}$. Furthermore, all Be components were stress relieved throughout the fabrication process to prevent the accumulation of stress.

Fig. 15

Draper labs creep test data.

The PMSA-180 requirement that a PMSA could survive launch with $<2.9\mathrm{nm}$ rms surface figure distortion was verified with a Webb flight segment. TRL6 was demonstrated by test. An unpolished Webb mirror segment B1 was assembled into a flight configuration PMSA and exposed to design limit loads with sine burst, random vibration, and acoustic testing. Its surface figure change as a function of each loading test was measured using a phase measuring electronic speckle pattern interferometer. The design limit load accelerations for every component within the PMSA were exceeded in each of these tests. Two acoustic tests were performed. The first test hard mounted the PMSA to a concrete wall. The second test suspended the PMSA for a “free–free” test (Fig. 16). The two different boundary conditions provide valuable information for finalizing the PMSA test environment. Neither test resulted in any measurable change to the PMSA surface figure.

Fig. 16

PMSA during first and second acoustic tests.

The tests showed that the surface figure was repeatable to within the noise floor of the metrology system, 14 nm rms (Figs. 17Fig. 18–19). Astigmatism and power figure terms were removed from the total surface change measurement because they can be compensated for on-orbit by adjusting the PMSA ROC or physical location via the cryogenic hexapod with the ROC actuator. The measurement of a surface figure error change that was smaller than the measurement noise floor was consistent with the pretest prediction. Nonlinear plastic material finite-element analysis predicted a surface figure deformation of only 1.6 nm rms.

Fig. 17

Figure change from exposure to three axis sine burst testing to design limit loads.

Fig. 18

Figure change from exposure to three axis sine burst testing to design limit loads and first acoustic test.

Fig. 19

Figure change from exposure to second acoustic test.

Although not affecting a determination of demonstrating TRL-6, there were two special circumstances associated with this test. First, the random vibration and acoustic levels were notched to maintain safe exposure levels on the PMSA, and a minor inconsistency was discovered with the design limit loads. A new test environment was defined and a minimal PMSA redesign performed to meet the new test environment. Second, although all flight PMSAs will be thermally cycled to 25 K before launch, for reasons of expedience and convenience, the B1 PMSA was only thermally cycled to 150 K. This was determined acceptable because the 150 K temperature subjected the mirror to $\sim 88$% of the beryllium cryo-strain and over 70% of adhesive mount strain. Additionally, an extensive qualification program was conducted for the bonded joints. Test samples were cycled 3 times between 15 and 383 K and subjected to static pull testing. These samples saw only a 12% reduction in ultimate strength following thermal cycling and still maintained a margin of safety of 7.4. This testing coupled with the 150 K that the B1 segment saw was more than sufficient to assure that the TRL6 vibration testing demonstrated the true robustness of the PMSA. TRL6 was achieved by demonstrating the technology to design a lightweight beryllium mirror to design limit loads, testing it to those loads, and showing surface figure stability after exposure to the design limit load, thus assuring that lightweight beryllium mirror technology could meet the Webb launch distortion requirements.